diff --git "a/designv11-34.json" "b/designv11-34.json" new file mode 100644--- /dev/null +++ "b/designv11-34.json" @@ -0,0 +1,12754 @@ +[ + { + "image_filename": "designv11_34_0002262_978-3-319-27333-4-Figure2.7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002262_978-3-319-27333-4-Figure2.7-1.png", + "caption": "Fig. 2.7 Orthogonally displaced journal bearing", + "texts": [ + " Elliptical journal bearings are slightly more stable toward the oil whip than the cylindrical bearings. In addition to this, elliptical journal bearing runs cooler than a cylindrical bearing because of the larger horizontal clearance for the same vertical clearance. Some other non-circular journal bearing confi gurations; Three-lobe journal bearing (symmetrical and asymmetrical, Fig. 2.5 ), Elliptical journal bearing (profi le is elliptic in cross-section, Fig. 2.6 ), and orthogonally displaced journal bearing (vertical offset, Fig. 2.7 ); have also been shown below. 2 Classifi cation of Non-circular Journal Bearings 7 e RL RL RL Fig. 2.5 Symmetrical three-lobe bearing In earlier works, the bearing performance parameters have been computed by solving the Reynolds equation only. Over the years, many researchers have proposed number of mathematical models. A more realistic thermohydrodynamic 2.3 Methods of Analysis 9 ( THD ) model for bearing analysis has been developed which treats the viscosity as a function of both the temperature and pressure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure6-1.png", + "caption": "FIGURE 6. STIFFNESS MEASUREMENT TEST FIXTURE SHOWING THE TEST SEAL AND THE DUMMY SEAL MOUNTED ON EITHER SIDE OF A ROTOR.", + "texts": [ + " Some of the T bristles in the front row of the T-F-T straight seal were fused. No fusion was observed on the tapered seals. The fusion was a result of a manufacturing process. The larger than predicted stiffness of the straight seals may be attributed to bristle fusion, making test data for the MF-T straight seal unreliable. In order to study the stiffness of brush seals when they are subjected to pressure loading, testing is performed in a highpressure chamber test rig. The test fixture [7], shown in Figure 6, consists of two large-diameter brush seals mounted on a circular base. The brush seal segment mounted on the right side is the test seal while the other one is an identical dummy brush seal. High-pressure air surrounds the entire fixture and lowpressure air is present inside the circular base, thereby creating a pressure drop across the bristles. The large-diameter rotor is simulated with a rectangular block with the rotor radius machined on either side. The rotor is mounted on a low-friction slide and can move radially relative to the test and dummy seal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure5.7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure5.7-1.png", + "caption": "Fig. 5.7 Gravity effect over an free-floating rigid mass", + "texts": [ + " Right: zyz convention. First an a-rotation around z, so that the neutral axis N (either the new x or y axes) is properly aligned. Second a b-rotation around the N axis. Finally a -rotation around the newest z-axis so that the frame reach the final orientation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198 xx List of Figures Fig. 5.6 Same cylinder with three different local reference frame assignments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243 Fig. 5.7 Gravity effect over an free-floating rigid mass . . . . . . . . . . . . . . 262 Fig. 6.1 Linear and angular velocities associated to geometric origin g of frame R1 and linear and angular velocities of the center of mass cm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 278 Fig. 8.3 Left: Swedish wheel, also known asMecanum wheel, invented by the Swedish engineer Bengt Ilon. Right: Omni-wheel, an upgraded design of a omnidirectional wheel . . . . . . . . . ", + "3 In ships and aircrafts, it is a convention to define inertial frame at the Earth surface with the vertical direction in the z-axis and positive through the center of the earth (such that the yaw angle coincide with a compass azimuth measure). Then the gravity field direction in the inertial frame coordinates would be given as: g(0) = \u239b \u239d gx0 gy0 gz0 \u239e \u23a0 = \u239b \u239d 0 0 g \u239e \u23a0 = gk For a rigid body, the gravity force acting on the center of mass can be calculated by the projection of the gravity acceleration over the moving frame as (Fig. 5.7) f (1) g = m g(1) = mRT g(0) (5.73) Remark Notice that for free-flying bodies there is no the gravity torque:ng = 0, since there is no anchor point from which the gravity force would produce any rotational motion, hence implying that the gravity force would not produce any moment. For non free-flying bodies, which suffer external forces applied at different points than the center of mass, this forces together with the force of gravity would induce a gravity torque. Then in a general way, the gravity torque of any body can be expressed to be 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002268_978-4-431-54361-9-Figure1.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002268_978-4-431-54361-9-Figure1.4-1.png", + "caption": "Fig. 1.4 The trajectory of a wavepacket for a skew scattering and b side jump. a After the scattering, the average trajectory of the electron is deflected by a spin-dependent angle \u03b4. \u03b4 is typically of order 10\u22122 rad [44]. b The center of the wave-packet is displaced laterally during the scattering. The lateral displacement \u0394y is about 10\u221211 m", + "texts": [ + " In this case, the \u03c3xy is reduced to \u03c3 2\u2212bands xy = e2 8\u03c0h \u222b dkzd\u03a9 f [ fF D(\u03b5\u2212(k)) \u2212 fF D(\u03b5+(k))], (1.18) where fF D(\u03b5) is the Fermi distribution function. Here d\u03a9 f is the f -space solid angle, which is the integral of the gauge field b( f ) = \u00b1 f | f |3 (1.19) due to the monopole at f = 0 over the infinitesimal surface in f -space corresponding to the small square dkx dky in k-space [10]. This gauge field strongly depends on k in the near degeneracy case, i.e. when f (k) is near the monopole. Let us consider an electron wavepacket approaching a central scattering potential (Fig. 1.4). Before scattering, the center of mass of the packet moves in a straight line thorough the periodic crystal. After scattering, the wavepacket is broken into a set of 1.2 Anomalous Hall Effect in Itinerant Ferromagnets 9 outgoing spherical waves. The motion of the center of mass again follows a straight line at a constant speed. If the electron spin s is normal to the plane of the illustration, spin-orbit interaction removes any symmetry between right and left. As a result, the trajectory after scattering may well differ from the one before scattering. For example, the new trajectory might be at an angle to the old one (Fig. 1.4a). This means that the electron acquires transverse momentum on scattering. The effect is called \u201cskew scattering\u201d. It can be derived from a classical Boltzmann equation if the differential cross-section has a left-right asymmetry. However, it vanishes in the first Born approximation, i.e. for weak scatterers. To obtain the skew-scattering contribution, we proceed to second order approximation. In the relation of \u03c1 A yx \u221d (\u03c1xx ) n , it gives n = 1. Skew scattering was proposed by Smit [40, 41]. A second possibility is that the new trajectory might be displaced by a finite distance from the old trajectory (Fig. 1.4b). This transverse displacement \u0394y is called a \u201cside jump\u201d [42, 43]. The \u0394y was estimated to be \u223c 10\u221211 \u221210\u221210 m, which may be large enough to yield AHE observed in ferromagnets. Surprisingly, the side-jump \u0394y per collision is found to be independent of the strength, range, or sign of the scattering potential. The side jump mechanism predicts n = 2 in \u03c1 A yx \u221d (\u03c1xx ) n . An intuitive picture of side-jump mechanism was presented in [44]. The original theory by Smit [40, 41] considered free electrons and short-range potentials" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003104_1.i010782-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003104_1.i010782-Figure5-1.png", + "caption": "Fig. 5 Desired motion of an agent on a target-centric formation.", + "texts": [ + " 3 and 4, it is observed that the desired location of UAV i in the target-centric formation is pdi xdi; ydi \u22a4 pT \u03b4di cos \u03d5i ; sin \u03d5i \u22a4 (42) where \u03b4di > 0 is the radial distance from T, and \u03d5i \u2208 \u2212\u03c0; \u03c0 is a reference angle formed at T by UAV i with respect to a certain reference axis. Next, the desired motion parameters for the targetcentric fixed formation are obtained from the desired formation geometry. In a target-centric fixed formation, \u03d5i is time invariant \u2200 i 1; 2; : : : ; N. A diagram of the desired motion of agent i is shown in Fig. 5a. From this figure, the desired speed vdi, heading \u03c8di, and angular speed \u03c9di of agent i are obtained as follows: vdi vT (43) \u03c8di \u03c8T (44) \u03c9di \u03c9T (45) In the following portion, the expressions of vdi;\u03c8di;\u03c9di are derived from the desired target\u2013agent geometry when each agent circumnavigates the target. To achieve target-centric circumnavigation, agents need to have a finite desired angular velocity \u03d6 with respect to the motion of the target. Therefore, in the case of circumnavigation, Eq. (42) is modified as pdi pT \u03b4di cos \u03d5i ;sin \u03d5i \u22a4; _\u03d5i \u03d6; \u2200 i 1;2;: : : ;N (46) Figure 5b depicts the desired motion of agent i while circumnavigating the target. The desired speed vdi, heading \u03c8di, and angular speed \u03c9di obtained from Fig. 5b are expressed as follows: vdi v2T \u03b42di\u03d6 2 2vT\u03b4di\u03d6 sin \u03c8T \u2212 \u03d5i 1\u22152 (47) \u03c8di tan\u22121 vT sin \u03c8T \u03b4di\u03d6 cos \u03d5i vT cos \u03c8T \u2212 \u03b4di\u03d6d sin \u03d5i (48) \u03c9di _\u03c8di 1 v2di \u2212aT\u03b4di\u03d6 cos \u03d5i \u2212 \u03c8T \u03c9Tv 2 T \u03b42di\u03d6 3 \u2212 \u03c9TvT\u03b4di\u03d6 sin \u03d5i \u2212 \u03c8T \u2212 \u03b4di\u03d6 2vT sin \u03d5i \u2212 \u03c8T (49) where aT \u2254 _vT is the translation acceleration of the target. Using the variables vdi;\u03c8di, and \u03c9di, the control laws are presented to achieve desired target-centric formations in the respective scenarios. To result in the desired target-centric formation in both scenarios, the control law must ensure the following conditions: 1) First, \u03d5i \u2212 \u03d5i\u22121 need to converge to the desired angular spacing \u2200 i 1; 2; : : : ; N", + " Thus, _\u03d5i converges to zero and \u03d6 in the case of target-centric fixed formation and circumnavigation, respectively. From Eq. (50), \u03c6i 2\u03c0 N \u21d2 mod2\u03c0 \u03d5i \u2212 \u03d5i\u22121 2\u03c0 N \u2212 \u03d1i D ow nl oa de d by U N IV E R SI T Y O F N E W S O U T H W A L E S on J ul y 12 , 2 02 0 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .I 01 07 82 Therefore, any interagent angular separation can be obtained by adjusting \u03d1i. Proof of the convergence was provided in Ref. [38]. Next, the tracking controller is presented, which drives the formation error to zero. Based on the desired formation geometry in Fig. 5, the formation errors are defined as follows: pei \u2254 xei; yei \u22a4 R \u03c8 i pdi \u2212 pi (52) \u03c8ei \u2254 \u03c8di \u2212 \u03c8 i (53) where R \u03c8 i cos\u03c8 i sin\u03c8 i \u2212 sin\u03c8 i cos\u03c8 i Using these errors, the formation control laws for both scenarios are presented as vi vdi K1i xei 1 x2ei y2ei p (54) \u03c9i \u03c9di K3i sin \u03c8ei\u22152 K2ivdi yei cos \u03c8ei\u22152 \u2212xei sin \u03c8ei\u22152 1 x2ei y2ei p (55) where inft\u22650 vdi > K1i > 0, and K2i; K3i > 0. The expression of vdi;\u03c8di;\u03c9di is given in Eqs. (43\u201345) and (47\u201349) for the target-centric fixed formation and circumnavigation, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001843_jahs.59.022006-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001843_jahs.59.022006-Figure2-1.png", + "caption": "Fig. 2. PAM TEF actuation system design concept.", + "texts": [ + " 2) and the \u201ctension\u2013torsion rod\u201d used in the SMART rotor design (Ref. 9). The TRF is a thin rod that is aligned spanwise and anchored on one end to the blade and on the other to the flap. It, therefore, provides a direct structural connection capable of transferring the considerable centrifugal (CF) loads on the flap into the blade structure without the need for thrust bearings. In addition, this device is designed to have minimal torsional stiffness, so that deflections of the flap may be achieved through elastic twisting of the TRF. Figure 2 shows a top-view design drawing of this actuation system embedded in a rotor blade. Note that the blade span has been shortened to allow for a greater detail of the inboard and outboard components. A quasistatic torque-balancing model of the flap system was developed, which allowed for the effects of different system parameters on flap deflections to be considered. This model balances the hinge moments required to deflect the flap against those available from the actuation system to determine the equilibrium flap deflection angles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001843_jahs.59.022006-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001843_jahs.59.022006-Figure5-1.png", + "caption": "Fig. 5. PAM volume testing.", + "texts": [ + " However, they also have significantly higher passive stiffness due to the braid filaments being more aligned with the pull axis. It was, therefore, not immediately clear which PAM yielded the best net system performance. In addition, the different PAM geometries can be seen to have significantly different internal air volumes. For the final characterization test, the PAMs were placed in a custombuilt volume test rig, which measures the change in volume of the PAM as a function of pressure and contraction ratio. This device, seen in Fig. 5(a), was mounted to an MTS machine to allow contraction ramps to be performed. By measuring the change in height of a column of water surrounding the PAM, the change in PAM volume could be calculated. The change in PAM volume is then added to the known initial internal air volume (calculated from the internal area of the bladder and PAM active length) to determine internal air volume, Vint as a function of pressure and contraction ratio. Figure 5(b) shows the results of this volume test for PAM 13. The behavior seen here is typical: Starting at \u03bb = 1.0, there is a large initial increase in volume with decreasing contraction ratio, but the volume levels off as the minimum contraction ratio is approached. Polynomial regressions were fit to the data to average out the effects of hysteresis. 022006-6 With the active and passive force profiles and the internal air volume for each PAM experimentally determined, the PAM selection process could begin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003088_tia.2020.2992579-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003088_tia.2020.2992579-Figure7-1.png", + "caption": "Fig. 7. Structure of proposed DSLSPMV machine.", + "texts": [ + " The squirrel cage of IM is embedded in the inner circumference of the rotor and the stator of IM is as the inner stator of DSLSPMV machine. The PMV machine and the IM share one rotor. The coupling of outer stator winding and the squirrel cage can be nearly eliminated when the rotor yoke length is large enough. Therefore, the starting performance and steady state performance can be decoupled, and the steady state performance is the same as the regular PMV machine. The wound winding also can be selected as the rotor starting unit when the operation braking toque is too heavy. The structure of DSLSPMV machine is shown in Fig.7. The PM pole-pair number Pe and outer stator winding polepair number PS satisfy: - S eP Z P (9) The inner stator, stator winding and rotor starting unit constitute totally an IM. Both squirrel cage and wound winding is selected as the rotor starting unit. If wound winding is adopted, the rotor wound winding pole-pair number Pr and inner stator winding pole-pair number PIM should satisfy: = r IMP P (10) If the synchronous speed of IM is much lower than synchronous speed of PMV machine, the rotor cannot be pulled into synchronization when the outer stator winding is connected to the power grid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001769_robio.2015.7418980-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001769_robio.2015.7418980-Figure2-1.png", + "caption": "Fig. 2 Pressing plate and laser sensors", + "texts": [], + "surrounding_texts": [ + "System (ADRS) has precluded their usage in aircraft assembly. Aiming at improving the accuracy of ADRS, three efforts are made in this paper. Firstly, a novel surface normal measurement and attitude adjusting method are proposed to guarantee the perpendicularity of holes. Secondly, an improvement compensation method is made for absolute positioning accuracy of industrial robot by using Gauss-Newton method. Thirdly, during the calibration of tool center point(TCP), which is one of the key technologies of offline programming control system, two balls are used in this method to solve the problem of low accuracy and inefficiency of traditional method using pointed end. Finally, both simulations and experiments are conducted to test the effect of the improvements and the results prove the improvements are effective.\nI. INTRODUCTION\nRiveting is the most popular connecting way in aircraft assembly. It is reported that 70% of accidents caused by fatigue failure happen at joints position and most fatigue cracks is from riveting holes[1]. Therefore quality of riveting holes have an extremely important effect on aircraft life. There are millions of riveting hole to drill, but the traditional manually drilling is not only low efficiency but also high labor-intensive. Even so, the quality of holes still cannot be guaranteed.\nIn recent years, industrial robot is applied in more and more factories. Production implementations of industrial robot in the aircraft manufacture and assembly are really active[2]. With the help of offline programming technology, industrial robot can be applied in more areas in factory.\nThe autonomous drilling robot system (ADRS), mentioned in this paper, is mainly composed of an industrial robot, end effector and control system, as shown in Fig.1.\nADRS can achieve high speed drilling automatically, but its accuracy is far from the requirements of aircraft assembly. Drilling perpendicularity, positioning accuracy and calibration of TCP all have great influence on ADRS accuracy[3].\nThe perpendicularity of the connecting holes has important effect on the quality of the riveting holes. The bad perpendicularity may lead to fatigue cracks, which could\n*Research supported by the National Natural Science Foundation of\nChina (No.61375085),\nM.Q. Lin, P.J. Yuan, Y.W. Liu, Q.C. Zhu and Y. Li are with School of Mechanical Engineering & Automation, Beihang University, China. (Email: winterchen2013@163.com, itr@buaa.edu.cn).\nH.J. Tan is with Chengdu Aircraft Industrial (Group) Co. Ltd, Chengdu,\nChina.\ndirectly affect the life of an aircraft[4]. The perpendicularity of robotic drilling depends on whether the normal at the drilling point coincides with the axis of the hole.\nIndustrial robot have high repeatability positioning accuracy (about 0.1mm), while the absolute positioning accuracy is far from being able to meet the requirements of automatic assembling in aviation manufacturing[5]. Therefore, the low accuracy of industrial robots has precluded their usage in many aerospace applications.\nCalibration of TCP is a key technology of offline programming control system. The traditional method to calibrate TCP is 4-point method. It requires controlling robot to make two pointed ends almost coincide. It is hard to guarantee the accuracy and low efficiency.\nIn order to improve the accuracy of ADRS, three efforts are\nmade on the above aspects.\nII. THE SURFACE-NORMAL MEASUREMENT METHOD AND\nMICRO-ADJUSTING MECHANISM\nA. Surface-normal Measurement\nThe perpendicularity of robotic drilling depends on whether the normal at the drilling point coincides with the axis of the hole. Thus, surface-normal measurement is key technology in drilling system. To get the attitude of the end effector, this paper presents a novel surface normal measurement by using three laser sensors.\nBefore the surface normal measurement process, the pressure plate is pushed against surface of workpiece. In the center of the pressure plate the drilling point P is marked. In this case, we can consider that the normal of surface coincides with the normal of the pressure plate. Therefore, the normal of surface can be acquired by measuring the normal of pressure plate.\n978-1-4673-9675-2/15/$31.00 \u00a9 2015 IEEE 1483", + "By using three laser sensors, the distance between the sensors and pressure plate of three points p1, p2 and p3 is obtained. Then with the structural parameters of end effector, the Cartesian coordinates of the three points can be expressed as follow.\n \n \n \n1 1 1 1\n2 1 1 2\n3 1 1 3\n, ,\n, ,\n, ,\nP x y z\nP x y z\nP x y z \n(1)\nWhere z1=L+K-a1, z2=L+K-a2, z3=L+K-a3. ai is the distance between the sensors and pressure plate when it is pushed out. L is the distance of pressure plate being pushed by cylinders, and it is obtained by linear position sensor. K is distance he sensors and pressure plate when it is not pushed out. According to the theorem that 3 points determine a plane, the equation of the pressure plate plane can be expressed as\n0Ax By Cz D (2)\nAs p1, p2 and p3 are all in the plane, we can obtain\n1 1 1 0Ax By Cz D (3)\n1 1 2 0Ax By Cz D (4)\n1 1 3 0Ax By Cz D (5)\nCombine (3), (4) and (5)\n1 3\n3 2\n1 2\n2\n2\n2\nz z D C\nz z A C\nd\nz z B C\ne\n \n(6)\nBy Combining (2) and (6) , we can obtain the equation of\nthe pressure plate plane.\n2 3 2 1\n1 3\n1 1\n2 0 z z z z\nx y z z z x y\n (7)\nThen we can obtain the normal n\n 2 3 2 1\n1 1\n, , ( , , 2) z z z z\nn A B C x y\n (8)\nMeanwhile, the equation of n can also be expressed as\n0 0 0x y z\nA B C\n (9)\nWhere 2 3\n1\nz\nx A\nz , 2 1\n1\nz\ny B\nz , C=2.\nB. Principle of drill attitude adjusting\nAdjusting attitude is the key technology in automatic drilling system[6][7]. As is shown in Fig.3, there is the principle of adjusting mechanism, which is composed of a spherical pair, a spherical pair bearing and two eccentric discs. The radii of two eccentric discs both are r, and the line O is the geometry axis of the big eccentric disc. While the line Ob is the eccentric axis of the big eccentric disc and it is also the geometry axis of the small eccentric disc. The eccentric axis of the small eccentric disc is the line Os. The drill centerline crosses the spherical pair and spherical plain bearing which is installed at the eccentric point of the small eccentric disc. The center of spherical pair coincides with the vertex of the drill, and that is point Osp. According to the target position, calculate the angles of each eccentric discs. As is shown in Fig.4, the adjusting area is a cone. The advantage of the adjusting mechanism is that the vertex of the drill is stationary, no matter how the drill centerline rotates, because the center of spherical pair coincides with the vertex of drill. The adjusting mechanism doesn\u2019t need to remove the vertex of drill to the drill point many times. It can be more efficient and save adjusting time for drilling.", + "III. COMPENSATION METHOD OF INDUSTRIAL ROBOT\nABSOLUTE POSITIONING ACCURACY\nCompare with the repeatability positioning accuracy of industrial robot (0.1mm), the absolute positioning accuracy (1-3mm) is far from meeting the requirements of automatic assembling in aviation manufacturing. When robot is moved by the tech pendant, the accuracy of positioning is determined by the repeatability positioning accuracy. But during the aircraft assembly, there are thousands of holes to drill. The efficiency of teaching machine model is too low, while Offline programming is efficient. But positioning accuracy of offline programming mainly depends on the absolute positioning accuracy. In order to solve this problem, this paper proposes a compensation method of industrial robot positioning accuracy. Firstly, the kinematic model of industrial robot is established by using the D-H modeling method[8], then analyze the forward kinematics and inverse kinematics. By using a laser tracker, the calibration model of robot was built with the Gauss-Newton method.\nA. Kinematic Model of Robot\nBefore the kinematics analysis, robot\u2019s kinematic model should be established. Industrial robots are generally link mechanism, which is composed of rigid links connected by joints. The main parameters of D-H model are showed in Fig. 6 including length of link , twist angle , setover , joint angle .\nAccording to the theory of rigid body translation and rotation, the transformation matrix between adjacent coordinates can be obtained.\n , 0,0, ,0,0 x,\n1 0 0 0 0 0\n0 1 0 0 0 0\n0 0 1 0 0 1\n0 0 0 1 0 0 0 1\n1 0 0 1 0 0 0\n0 1 0 0 0 0 \u00a0\n0 0 1 0 0 0\n0 0 0 1 0 0 0 1\ni i i i i\ni i\ni i\ni i\ni\ni i\ni i\ni i i i i i i\ni\nA Rot z Trans d Trans a Rot\nc s\ns c\nd d\na\nc s\ns c\nc s c s s a c\ns c\n \n \n \n \n \n \n\n\n \n \n\n 0\n0 0 0 1\ni i i i i\ni i i\nc c s s\ns c d\n \n \n \n(10)\nWhere \u00a0 , c is cos s is sin .\nB. Forward Kinematics of Robot\nWith the structure parameters and joints rotational angle, to calculate the pose of end of robot in the base coordinates, namely, forward kinematics.\nSix translation matrixes are obtained by substituting the D-H model parameters of robot into eq.(10). Then the translation matrix between the base coordinates and end of robot is obtained.\n6 1 2 3 4 5 6 \u00a0\n0 0 0 1\nx x x x\ny y y y\nz z z z\nn o a p\nn o a p A A A A A A\nn o a p T \n(11)\nWhere\n 1 4 6 4 5 6 1 23 4 6 5 6 23 5 6xn s c s s c c c c s s c c s s c \n 1 23 4 6 5 6 23 5 6 1 4 6 4 5 6yn s c s s c c s s c c c s s c c \n 23 5 6 23 4 6 4 5 6zn c s c s s s c c c \n 1 4 6 4 5 6 1 23 4 6 4 5 6 23 5 6xo s c c s c c c c s c c c c s s c \n 1 23 4 6 4 5 6 23 5 6 1 4 6 4 5 6yo s c s c c c s s s s c c s s c s \n 23 5 6 23 4 6 4 5 6zo c s s s s s c c s \n 1 4 5 1 23 4 5 23 5xa s s s c c c s s c \n 1 23 4 5 23 5 1 4 5ya s c c s s c c s s " + ] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure6.22-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure6.22-1.png", + "caption": "Figure 6.22 A spatial cam mechanism that allows rolling without slipping.", + "texts": [ + "377) \ud835\udf062 12zteu\u03032\ud835\udf0301 (\u0302I + \ud835\udf0612B\u0302)\u22121B\u03022(\u0302I + \ud835\udf0612B\u0302)\u22121e\u2212u\u03032\ud835\udf0301 z = r2 2 (6.378) Referring to Eq. (6.363), note that z = z(\ud835\udf0301, \ud835\udf0302). Therefore, Eqs. (6.377) and (6.378) contain two unknowns, which are \ud835\udf0612 and \ud835\udf0302. Note also that Eqs. (6.377) and (6.378) are highly nonlinear and therefore they can be solved only by using a suitable iterative numerical method. However, after \ud835\udf0612 and \ud835\udf0302 are thus determined, the coordinates of the contact points can be found readily by using Eqs. (6.375) and (6.376). (a) Kinematic Description of the System Figure 6.22 shows a modified version of the spatial cam mechanism considered in the previous example. The modification is such that the previous revolute joint 02 is converted into a cylindrical joint and the previous link 2 is divided into two links 2 and 3, which are connected by a revolute joint 23. Upon this division, the spherical cam is attached to 3 and it forms the cam joint 13 together with the elliptical cam, which is still attached to 1. Thus, the mobility of the modified mechanism becomes \ud835\udf07 = 3 as long as there is no sticking friction between1 and3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003619_s11668-020-01031-4-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003619_s11668-020-01031-4-Figure8-1.png", + "caption": "Fig. 8 Distribution of frictional heat generation (a) linear pressure and (b) constant pressure", + "texts": [], + "surrounding_texts": [ + "In the previous section, it was presented the full details of the developed numerical solution of the thermal-structural problem of the friction clutch system to investigate the effect of the variation of applied pressure during the sliding period (heating phase) on the frictional heat generated, contact pressure and temperature. The first function of applied pressure was constant with time (P= 0.5 MPa), while the second one is increased linearly with time from P = 0 at (ts= 0) to the final value (P = 0.5 MPa) at the end of the sliding period at (ts= 0.8). The slipping times for both cases were calculated based on the equations of motion [19], where the slipping times to reach the full engagement phase (when all parts rotate together at the same speed without slipping between each other) of each case are 0.4 and 0.8 s corresponding to the constant and linear applied pressure functions, respectively. Figure 4 illustrates the distribution of the contact pressure with disc radius during a single engagement applying the constant and linear functions of applied pressure. The highest values of pressure in contact are found as P = 1.44 MPa and P = 1.10 MPa (at the pressure plate side) corresponding to the constant and linearly functions of applied pressure, respectively. The maximum contact pressure was reduced approximately 24% when applied linear function instead of the constant function of pressure. Figure 5 shows the distributions of the contact pressure during the sliding period when applying the linear and constant functions of pressure. These results illustrated that the contact area was increased by 12% when applying the linear function of pressure instead of the constant one. The reason for such results is gradual increasing in applied pressure according to the applied the linear function. While, when applied the constant pressure function, the significant stresses and deformations were appeared in the contacting surfaces of clutch system. Figure 6 exhibits the variation of maximum surface temperature during the sliding period when applying the constant and linearly functions of applied pressure. The highest value of surface temperature is found to be T = 382.39 K that occurred at r = 0.0678 m and t = 0.6ts (pressure plate side) when the applied a constant pressure function. However, the maximum value of surface temperature for linear function of applied pressure is T = 350.03 K that located at r = 0.0678 m and t = 0.7ts on the pressure plate side. Figure 7 shows that the surface temperatures of the friction clutch disc under different working conditions during the whole engagement period (applying linear and constant pressures). The convenience of applying linear acting pressure function reflects on the Fig. 4 Maximum contact pressure during the sliding period Fig. 5 Distribution of contact pressure (a) linear pressure and (b) constant pressure temperature that occurred at the contact surfaces. As the pressure changes linearly from 0 to 0.5 MPa, and at the same time there is a gradual equalization of angular speed of drive and driven part. This will lead to appear temperatures on contact surfaces are lower than the temperatures that occurred in the case of constant pressure. By observing Figs. 5 and 7, it can be concluded that the value of pressure, which occurred on contact surfaces, effects on the value of the generated temperature in the assembly. Therefore, it can be concluded that when the function of the acting pressure of the frictional clutch is linear, this will lead to increase the lifetime of the contacting surfaces of friction clutch. Figures 8 exhibits the variation of the heat generated on the friction clutch disc under different working conditions during the whole engagement period (applying linear and constant pressures). These results proved that any change in applied pressure function considerably affects theFig. 6 Maximum surface temperature during the sliding period thermoelastic behavior of the friction clutch assembly. When applying the constant pressure, the maximum heat generated is approximately 3.75 MJ. While, when applying the linear function for the applied pressure, the maximum heat generated is approximately 1.2 MJ on the pressure plate side. Generally, the heat generated decreases gradually until reach zero at the end of the sliding period, where there is no relative motion anymore between the contact parts of the friction clutch assembly." + ] + }, + { + "image_filename": "designv11_34_0001741_eleco.2015.7394513-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001741_eleco.2015.7394513-Figure2-1.png", + "caption": "Fig. 2. Reference frames, rotation angles and rotation rates", + "texts": [ + " In this work, the dynamical model for a plus-type quadrotor is given. After that the Crazyflie quadrotor is flown and realtime data is collected. Then this data is analyzed using Matlab to find an appropriate ARX (auto-regressive exogenous) model. 2. Dynamical Model of Quadrotor A quadrotor model can be evaluated in two parts: the motor and the body dynamics. The body dynamics can be divided into another two: rotational and translational dynamics. See fig. 1. Before giving the model equations, the coordinate frames and Euler angles have to be described. See fig. 2. The position and orientation of the quadrotor will be given relative to a fixed coordinate frame, which is called the inertial frame. The X and Y axes of the frame are placed parallel and coincident to the ground while Z axis is pointing upwards in a right handed configuration. This configuration is also described by East, North, Up (ENU) coordinates. (1) The mobile frame placed to the center of gravity of the quadrotor is called the body frame. The XB axis points the forward direction of the quadrotor, the YB axis points to the left and the ZB axis points up in a right-handed configuration. The most important problem in quadrotor control is its orientation in space. The well-known Euler angles representation is suitable for this purpose. The Euler angles (roll), (pitch) and (yaw) are the rotation angles about the axes X, Y and Z respectively (see fig 2). (2) The yaw-pitch-roll (YPR or ZYX) composite rotation matrix [2] that transforms an orientation from the body frame to the inertial frame is given below. The angular velocities (attitude rates) of and r are shown in fig. 2. All motor on the quadrotor contribute to relative to their angular velocities. It is pos motor produces the force , which is square of the angular speed and the total thr individual thrusts. This is given below constant and is the speed of motor Mi. ! \"# $ %% & $ %% ' ! & The torques that act about the roll and given below. () * + ! (, + - The torque about the yaw axis is diffe above, and expressed as follows where b is t (. / + -! 0 !! + ! 0 (# $()(,(.& 1 *! + ! ! + !/ + -! 0 !! + ! This work only covers the rotational dy due to the lack of precise motion capture sys It can be said that a traditional quadrotor for XB, YB and ZB axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003114_042079-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003114_042079-Figure5-1.png", + "caption": "Figure 5. The process of overcoming obstacles: (a) \u2013 phase of entry, (b) \u2013 phase of descent.", + "texts": [ + "9 m/s), which corresponds to the second lower gear of the tractor MTZ-82.1. It was set by applying independent virtual engines to each wheel. The engine speed was 10.9 for the rear wheels and 24.4 rpm for the front wheels. More detailed simulation parameters are presented in table 1. ICMSIT 2020 Journal of Physics: Conference Series 1515 (2020) 042079 IOP Publishing doi:10.1088/1742-6596/1515/4/042079 \u2022 linear speed of the machine-tractor unit. Figure 4 shows the simulation process with real-time display of parameters. Figure 5 shows the process of overcoming various obstacles with a machine-tractor unit. Overcoming a single linear obstacle in all series of experiments occurred without dangerous rolls and oscillations of the tractor. Overcoming a single sequential obstacle caused insignificant tractor oscillations. At the time of moving the obstacles with the front wheels, the balanced suspension of the front axle was actively working, reducing lateral vibrations. When overcoming obstacles with the rear wheels, the oscillations were more significant, due to their rigid attachment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000762_gt2013-95442-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000762_gt2013-95442-Figure3-1.png", + "caption": "FIGURE 3. Front-drawing of the steam test rig", + "texts": [ + " The axial position of the rotor is adjustable by using an integrated displacement device (6). For the right flow side, called \"motor side\", a by pass (7) is installed to regulate the pressure drop at the last two brush seals. The cross-section shows an installation of two different seal arrangements tested in the last years. Named by the location near to the motor\u00b4s (motor side, MS) and the displacement\u00b4s (displacement side, DS) end, the arrangements are labelled with no. 9 (MS) and 10 (DS). Furthermore the high pressure chamber and the chamber covers are named with 8 and 11. FIGURE 3 shows the test rig in a frontal drawing. In order to control and to fix the centric position between the rotor and the high pressure casing, it is possible to heat (14) or to cool (13) the columns (12). A non-uniform temperature distribution to the columns is caused by the non-constant operation of the power plant, the insulation and the discharge quantity of the rig. A shut off of the cooling results in a distension of the columns and produces an eccentric rotor position. As described, the test rig is located in a power plant in Braunschweig, similar to a normal turbine it is possible to control the inlet pressure by a control valve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003599_j.compositesb.2020.108452-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003599_j.compositesb.2020.108452-Figure7-1.png", + "caption": "Fig. 7. Test specimen types.", + "texts": [ + " Application of film contraction during curing and manual compression using roller has reduced the bridging effect on the intersection area. The compression applied using the heat shrinkable tape not only reduces the thickness difference by removing excess resin but also contributes to the minimization of the fiber bending deformation. This contributes to the minimization of the resin content at resin rich areas and, thus, strength improvement can be expected. Compressive strength test was conducted for the helical rib. For the testing, 2 types of specimens were prepared according to the areas of interest as shown in Fig. 7. \u201cSpecimen 1\u201d is extracted from IH region for testing of the IH strength and \u201cspecimen 2\u201d is extracted from PH region for testing of the PH strength. For the compression testing, alignment is important. However, since the extracted specimens have curvature, it was difficult to attach tabs on the specimen ends and to maintain specimen alignment along the applied load direction. Thus, a hollow jig was manufactured for potting of the helical rib inside the hollow cavity. Here, the adhesive EA9394 was used for the potting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003442_isitia49792.2020.9163776-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003442_isitia49792.2020.9163776-Figure2-1.png", + "caption": "Fig. 2. Parameter of AIV in Cartesian Coordinate Systems", + "texts": [ + " These wheel rotation speeds generate linear and angular velocity by using forward kinematics as follows: = cos 0sin 00 1 2 2\u2212 (3) Given certain linear and angular velocity = , wheel rotation speeds can be obtained by applying inverse kinematics as follows: = 1 1 21 \u2212 2 (4) Actual position and orientation of robot can be updated by applying Eulerian integration to its velocity = = as follows: ( + 1)( + 1)( + 1) = ( )( )( ) + ( )( )( ) (5) III. PROPOSED INDOOR NAVIGATION Our proposed indoor navigation is designed with respect to kinematics of AIV. Structure of indoor navigation is displayed in Fig. 2. Indoor navigation requires five input that are initial position of AIV , target position , map, distance of LIDAR sensor , and AIV position of odometry . It results two output that are linear velocity and angular velocity for AIV mechanism. Based on these velocity commands , AIV mechanism moves by wheel rotation speed by implementing (4). AIV movements are read by odometry to produce by applying (3) and (5). Details of indoor navigation is depicted in Fig. 3. AIV navigation consists of global and local path planner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure7.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure7.6-1.png", + "caption": "Figure 7.6 D\u2013H convention for a gripper.", + "texts": [ + "39) \ud835\udf03m = \u2222[u\u20d7(m\u22121) 1 \u2192 u\u20d7(m) 1 ] about u\u20d7(m) 3 (7.40) \ud835\udefde = \u2222[u\u20d7(m) 3 \u2192 u\u20d7(e) 3 ] about u\u20d7(m) 1 (7.41) \ud835\udeffe = \u2222[u\u20d7(m) 1 \u2192 u\u20d7(e) 1 ] about u\u20d7(e) 3 (7.42) However, in a practicable manipulator, for the sake of simplicity, the end-effector is preferably shaped and affixed to the last link so that it becomes possible to take e(Oe) coincident with m(Om). Such an end-effector can be designated as a regular end-effector. The most typical example of a regular end-effector is a gripper with two fingers, which is illustrated in Figure 7.6. The relevant definitions concerning the gripper shown in Figure 7.6 are given below. P = Om: Tip point (origin of the last link frame m) (7.43) R = Om\u22121: Wrist point (origin of the penultimate link frame m\u22121) (7.44) u\u20d7(m) 3 = u\u20d7a: Approach vector (7.45) u\u20d7(m) 1 = u\u20d7n: Normal vector (7.46) u\u20d7(m) 2 = u\u20d7s: Side vector (7.47) In the preceding definitions, u\u20d7n is called a normal vector because it is perpendicular to the gripper plane formed by the fingers. Similarly, u\u20d7s is called a side vector because it indicates the lateral direction in the gripper plane. The reason why u\u20d7a is called an approach vector is due to imagining that the gripper approaches the object to be gripped essentially in the direction of u\u20d7a", + " The main points of the compact and detailed formulations are explained in the following sections. For the sake of avoiding the complications that do not normally occur in the practicable cases, these formulations are obtained here by assuming that the following simplifying features exist for the first joint and the end-effector, which are explained in Chapter 7. \u2217 The base frame is positioned so that \ud835\udefe01 = 0, h01 = 0, and b01 = 0. Moreover, in most of the cases, it may be oriented so that \ud835\udefd01 = \ud835\udefd1 = 0, too, as illustrated in Figure 7.4. \u2217 The end-effector is regular as illustrated in Figure 7.6. That is, e(Oe) = m(Om). C\u0302(k\u22121,k) = eu\u03031\ud835\udefdk eu\u03033\ud835\udf03k (8.1) r\u20d7k\u22121,k = r\u20d7Ok\u22121Ok = bk\u22121u\u20d7(k\u22121) 1 + sku\u20d7(k) 3 (8.2) r(k\u22121) k\u22121,k = u1bk\u22121 \u2212 u2sks\ud835\udefdk + u3skc\ud835\udefdk (8.3) In the first stage of the compact formulation, Eqs. (8.1) and (8.3) are combined as shown below to obtain the homogeneous transformation matrix between the links k\u22121 and k . H\u0302 (k\u22121,k) = \u23a1 \u23a2\u23a2\u23a3 C\u0302(k\u22121,k) r(k\u22121) k\u22121,k 0 t 1 \u23a4 \u23a5\u23a5\u23a6 (8.4) As for the element-by-element expression, H\u0302 (k\u22121,k) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 c\ud835\udf03k \u2212s\ud835\udf03k 0 bk\u22121 c\ud835\udefdks\ud835\udf03k c\ud835\udefdkc\ud835\udf03k \u2212s\ud835\udefdk \u2212sks\ud835\udefdk s\ud835\udefdks\ud835\udf03k s\ud835\udefdkc\ud835\udf03k c\ud835\udefdk skc\ud835\udefdk 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 (8", + "17) by using the following equalities. C\u0302ku1 = ?\u0302?keu\u03031\ud835\udefek u1 = ?\u0302?ku1 (8.19) C\u0302ku3 = ?\u0302?keu\u03031\ud835\udefek u3 = ?\u0302?knk = ?\u0302?k\u22121nk (8.20) As for the end-effector, it is convenient to use the following definitions. C\u0302 = C\u0302m = C\u0302(0,m) \u2236 Orientation matrix of the end-effector, i.e.m (8.21) p = rP = rm = r(0)0,m \u2236 Location matrix of the tip point P = Om (8.22) r = rR = rm\u22121 = r(0)0,m\u22121 \u2236 Location matrix of the wrist point R = Om\u22121 (8.23) According to the assumption that the end-effector is regular such as a gripper shown in Figure 7.6, the column matrices r and p are related to each other as follows: r(0)OP = r(0)OR + r(0)RP = r(0)0,m + dmC\u0302(0,m)u(m) a = r(0)0,m + dmC\u0302(0,m)u(m\u2215m) 3 \u21d2 p = r + dmC\u0302u3 (8.24) In Eq. (8.24), dm is the length of the end-effector between the wrist and tip points. Since it is a length measured along u\u20d7(m) 3 , it is conceived as an offset and called a tip point offset. Position and Motion Analyses of Generic Serial Manipulators 205 By using the previous orientation and location equations derived for a link k , the following equations can be written for the position matrices C\u0302 and r of the end-effector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000881_elk-1212-20-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000881_elk-1212-20-Figure2-1.png", + "caption": "Figure 2. Block diagram and main subsystems of the MS&H simulator.", + "texts": [ + " On the other hand, the second type of simulator is the most realistic, but it is complex and expensive because of using mechanical platforms. This motivates us to choose a different kind of RTHIL simulator, the so-called mixed software and hardware (MS&H) simulator. This kind of simulator moderately enjoys the advantages of both previous types, i.e. it is sufficiently simple to construct and accurate for RTHIL verification in many applications. The block diagram and main subsystems of the MS&H simulator for the satellite ACS are shown in Figure 2. The main components of the MS&H simulator are as follows: a. Digital computer for modeling the satellite\u2019s attitude motion and environmental effects (orbit, gravity, magnetic field, etc.). b. Attitude determination and control hardware, for implementation of attitude control laws, command and data handling (C&DH) procedures, and synchronizing of the software and hardware units. c. Peripheral component interconnect (PCI) interface, which interconnects the ACS (hardware) and digital computer (software) in a real-time closed loop structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003675_ecce44975.2020.9236229-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003675_ecce44975.2020.9236229-Figure4-1.png", + "caption": "Fig. 4: (Left) The hexagonal coil used in FE model. (Right) The hexagonal coil model used in analytic calculation", + "texts": [ + " 3D FE RESULTS AND VALIDATION In order to validate the proposed analytic approach for inductance calculation, a 3D FE model of a three-phase, six-pole air-cored induction machine was built in COMSOL. Singlelayer fully-pitched coil windings were considered for the stator and rotor, with qs = 3 and qr = 2 coils per pole per phase for the stator and rotor respectively. Since simple rectangular or hexagonal coil shapes would cause end-coil intersections in the full winding\u2014which cannot be tolerated in a 3D FE model\u2014 hexagonal coils with end-steps were adopted for the 3D FE model, as shown in Fig. 4. The coil dimensions are given in Table VI and two snapshots of the 3D FE winding models are shown in Fig. 5. The selected end-step hexagon coil shape is a reasonable approximation of flat hexagonal coils adopted in the real air-cored induction machine prototype shown in Fig. 6, which was built to allow future experimental validation. However, if necessary, the proposed analytic approach can handle more complicated and refined shapes. The analytic and FE results are compared in Table VII, where both results are calculated when stator phase A and 5818 Authorized licensed use limited to: Univ of Calif Santa Barbara" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003402_s11071-020-05852-8-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003402_s11071-020-05852-8-Figure10-1.png", + "caption": "Fig. 10 Angle correction of the measured data to convert acceleration from normal tangential to Euclidean", + "texts": [ + " The acceleration transducers are tightly glued to the mass on the tip of the beam, for recording the acceleration of the motion in both directions of the trajectory. Because the actual motion of the mass travels on a curved trajectory, and the transducers used are single direction, it is not possible to directly record measurements in the rectilinear directions (u and v), rather accelerations are recorded in normal and tangential coordinates, which are later corrected with the appropriate angle to horizontal and vertical components (Fig. 10). It should be mentioned that although only the horizontal component (x\u0308) is being compared in the present study, both components are collected for completeness. Post-processing of the collected data includes a step of two-way (ascending and descending direction of the data set) low-pass filtering to eliminate high-frequency noise, and minimize phase shifts to the signals. The filter used is a Butterworth low-pass filter of order 8 and a cutoff frequency of 12Hz. The rigid surface is interchangeable to different sets, depending on the case of interest" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001436_0954406214523581-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001436_0954406214523581-Figure2-1.png", + "caption": "Figure 2. Calculation coordinate for the gas journal bearing with three-axial grooves.", + "texts": [ + "26 Due to its rapid convergence rate and minimal calculation error, many scholars have studied DTM and applied it to engineering.27\u201329 In order to reduce the calculation cost of solving the gas film forces of the self-acting gas bearing with three-axial grooves by the iterative method and improve the calculation accuracy, the DTM is employed to solve nonlinear pressure distribution of the gas film of the bearing. The calculation coordinate for a gas journal bearing with three-axial grooves is shown in Figure 2. Ob is the center of the bearing, Oj is the center of the journal, is the groove width angle, is the arc bushing angle of each bearing bush, is the location angle of the pad which is calculated from negative y axis to the leading edge of the first pad, \u2019 is the dimensionless circumferential coordinate of the bearing which starts from deviation line, f is the dimensionless angle which is calculated from negative y axis to gas film location, is the dimensionless deviation angle, R is the radius of bearing journal, h is the thickness of gas film, fx and fy are the nonlinear gas film forces in the negative directions of x and y axes, respectively, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.35-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.35-1.png", + "caption": "Figure 9.35 Elbow-down and elbow-up poses of the manipulator.", + "texts": [ + " The left shouldered and right shouldered poses of the manipulator that correspond to \ud835\udf0e1 = + 1 and \ud835\udf0e1 = \u2212 1 are shown in Figure 9.34. They are almost the same as those of a regular Puma manipulator with the difference that the wrist point R is replaced with the tip point P. (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e3 that arises in the process of finding \ud835\udf033 by using Eqs. (9.584)\u2013(9.586). The elbow-down and elbow-up poses of the manipulator that correspond to \ud835\udf0e3 = + 1 and \ud835\udf0e3 = \u2212 1 are shown in Figure 9.35. They are similar to those of a regular Puma manipulator with the difference that both the wrist and tip points are involved in the formation of these poses. In this case, the manipulator can have only one kind of multiplicity, which is the same as the second kind of multiplicity mentioned above in Section 9.7.4.1. In this case, the manipulator may have two kinds of position singularities, which are described and discussed below. (a) First Kind of Position Singularity Equation (9.575) implies that the first kind of position singularity occurs if the position of the end-effector is specified so that p1 = p2 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001404_dscc2013-3941-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001404_dscc2013-3941-Figure2-1.png", + "caption": "Figure 2. THE CONFIGURATION OF THE LANDFISH.", + "texts": [ + " Both systems have a single controlled degree of 1 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use freedom, the manipulation of which can regulate the systems\u2019 heading and generate forward locomotion. The control objective was to achieve steady translation in a specified direction. In both cases, single input proportional feedback control was shown to be effective in stabilizing the systems toward a specified heading, a simple result despite the nonlinearities of such systems. Figure 2 depicts the landfish schematically with its parameters. Denote the center of mass of the tail link T , and that of the head link H. The end of the head link is connected via a hinge to the tail link at T . The distance from the nonholonomic back wheel\u2019s ground contact point to T is a, while the distance from T to H is b. The head is supported by an unconstrained caster. The symbols mt , It , mh, and Ih represent the masses of the tail and head links, respectively, and their moments of inertia relative to vertical axes passing through their centers of mass" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003437_s00170-020-05886-7-Figure24-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003437_s00170-020-05886-7-Figure24-1.png", + "caption": "Fig. 24 A single-degree-of-freedom translational flexure stage made of 1-mm-thick fused silica [137]", + "texts": [ + " The amorphous phase of SiO2, fused silica (a-SiO2), is one of the most common high-quality optical substrates transparent to a broad spectrum (from about 170 nm to above 2.5 \u03bcm). It is inert to most chemicals, which also makes it a wide range of material selection for broad biological studies. Fused silica has a very low coefficient of thermal expansion (0.55 10\u20136/ \u00b0C) and is resistant to thermal shocks. It also possesses a low density (2200 kg/m3) and Young\u2019s modulus (72 GPa) comparable to that of aluminum alloys. It is an attractive material for flexures, though initially counterintuitive (Fig. 24). The AM technology for precision engineering motion devices has been first reviewed in academic society. The AM fabrication method with its materials and properties are compared with those current traditional manufacturing methods. Moreover, the common defects and its proper inspection methods in AM technology which directly influence the performance of nanopositioning systems are discussed, where fatigue is the most sensitive property. Some researches focused on thermal, static, and dynamic performance; new material and and structure design for further applications are also summarized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure2.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure2.3-1.png", + "caption": "Fig. 2.3 Multi-particle systems. Left: a single rigid body B with associated reference frames and main points position vectors. Right: the classical 4 bar linkage (with a virtual bar 1 being always static)", + "texts": [ + "16 Type I area D for Green\u2019s Theorem with boundaries given by 4 curves C1, C2, C3, and C4 . . . . . . . . . . . . . . . . . . . . . . . . . 96 Fig. 1.17 Equivalent 2D region is s-t coordinates of the Kelvin\u2013Stokes Theorem. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 99 Fig. 2.1 Reference frame and the particle position . . . . . . . . . . . . . . . . . 102 xix Fig. 2.2 2 different paths for the same particle to move from point 1 to point 2 . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 104 Fig. 2.3 Multi-particle systems. Left: a single rigid body B with associated reference frames and main points position vectors. Right: the classical 4 bar linkage (with a virtual bar 1 being always static) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 113 Fig. 2.4 Real path y\u00f0x\u00de and -varied path y\u00f0x; \u00de . . . . . . . . . . . . . . . . . . . 129 Fig. 2.5 Spherical surface segment defined as x2 \u00fe y2 \u00fe z2 \u00bc 1 in the positive 3D quadrant. Left: Restriction defined as / \u00bc x2 \u00fe y2 \u00fe z2 1 \u00bc 0 gives outwards gradients", + " qn \u239e \u239f\u23a0 \u2208 R n is called a set of generalized coordinates of a system, if and only if the number n of its elements is necessary and sufficient (equivalently to a minimal number of linearly independent variables) to define the configuration or the positional status of the system uniquely, (Banerjee 2005). The dimension n of such coordinates is given by the rule (Meirovitch 1970): n = 3N \u2212 h for a given N particles and h holonomic1 constraints. Important Issues Arising from the Generalized Coordinates Concept: \u2022 The set q may not be unique since for the same system many combinations may fulfill requirements in Definition 2.9. Example 2.1 The four bar linkage in Fig. 2.3b can be characterized uniquely with only one generalized coordinate either absolute angles \u03b82, \u03b83, or \u03b84, or relative angles \u03b85, or \u03b86. 1Holonomicmeans integrable in the sense of Frobenius in the framework of differential geometry. In a practical situation it can be understood as the ability of a constraint expression in a differential form: axdx + aydy + azdz + a0dt = 0 to bewritten as an algebraic restriction in the form f (x, y, z, t) = 0. If the differential form cannot be written as an equality and in terms of the non-differentiated variables x, y, z, it said to be non-holonomic. Dependence of time give rise to the terminology scleronomic, when it does not depend explicitly on time, and otherwise rheonomic, (Meirovitch 1970). 2.2 Lagrange Mechanics for Multi-particle Systems 113 Remark The rigid body example of Fig. 2.3a needs particular attention and is the central core of this text. However it needs 6 generalized coordinates to establish its position (3 independent variables) and attitude (another 3 independent variables). \u2022 The spacial coordinates (position) of every and each of the N particles of the system can be described with these generalized coordinates: d1 = d1(q1, q2, q3, . . . , qn, t) ... dN = dN (q1, q2, q3, . . . , qn, t) equivalently d j = d j (q, t) \u2208 R 3 For example: Each particle in the four bar linkage in Fig. 2.3 can be expressed in terms of the unique chosen variable; either \u03b82, \u03b83, \u03b84, \u03b85, or \u03b86. \u2022 The velocity of any particle can be written as v j = d\u0307 j (q, q\u0307) = n\u2211 i=1 \u2202d j \u2202qi q\u0307i + \u2202d j \u2202t = \u2202d j \u2202q q\u0307 + \u2202d j \u2202t (2.24) The term \u2202d j \u2202qi represents the tangent direction of d j w.r.t the coordinate qi , in a different (often larger) finite n-dimension space defined by q \u2208 R n; and the vector 114 2 Classical Mechanics q\u0307 = \u239b \u239c\u239d q\u03071 ... q\u0307n \u239e \u239f\u23a0 \u2208 R n is the generalized velocity. The last term \u2202d j \u2202t in (2", + " On the other hand, the linear velocity depends on the cross product of the angular velocity and a position vector, making that the linear velocity is valid only along the axis defined by the position rc, making it to be a line vector. \u2022 Notice also that the acceleration of the center of mass can be computed from (3.81) replacing the distance to any particle for the distance of the center of mass: d\u0308c = R ( v\u0307 + \u03c9\u0307 \u00d7 rc + \u03c9 \u00d7 v + \u03c9 \u00d7 (\u03c9 \u00d7 rc) ) 5.2 The Kinetic Energy 239 Fig. 5.4 Linear and angular velocities associated to the center of mass and on point p 5.2 The Kinetic Energy The kinetic energy of the body B of Fig. 2.3 can be easily found as the addition of the small amount of kinetic energy of each particle (Fig. 5.4): K = 1 2 \u222b B \u2225\u2225d\u0307p \u2225\u22252 dm (5.16) Remark Since the euclidean norm of a vector is its magnitude, i.e.: \u2016a\u20162 = \u2016a\u2016 \u221a a2 x + a2 y + a2 z with the followingproperty:\u2016a\u20162 = a\u00b7a = aTa, it is of no importance whose frame coordinates are used to compute de velocity of every particle. In other words, the kinetic energy does not depend on the choice of any particular reference frame. The position of any particle dp can be given by the rigid motion expression (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003970_ssrr50563.2020.9292576-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003970_ssrr50563.2020.9292576-Figure4-1.png", + "caption": "Fig. 4. Mechanical elements of the track", + "texts": [ + " However, shapes in which the track can be transformed are limited, and additional actuators are required, leading to an increase in weight. III. MECHANICAL DESIGN In contrast to all alternatives mentioned in the previous section, the MW-track is highly flexible and can adapt to rough terrains. As a result, high mobility is realized via a simple 150 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 04,2021 at 07:32:03 UTC from IEEE Xplore. Restrictions apply. structure. The MW-Track is composed of a metal chain belt and leaf springs, as indicated in Fig. 4 and described as follows. The track is composed of leaf springs that are attached to a metal chain, which is partially wrapped around a sprocket and supported by a holder. The leaf springs have uniformly spaced holes and are fixed to the attachment using screws. Furthermore, a spacer is located between these screws. The leaf spring passes through the center of rotation of the chain, enabling it to easily bend in both directions. Similar to wheels, a large loop facilitates easy movement past obstacles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001576_ls.1326-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001576_ls.1326-Figure1-1.png", + "caption": "Figure 1. Calculation coordinate for a TPFPAJB.", + "texts": [ + "14\u201317 Focusing on micro-grooved three-pad fixing pad aerodynamic journal bearings (TPFPAJBs), this study systematically evaluates the effect of the groove geometry on the load performance by the multi-grid finite element method. Special attention is given to the effect of the depth, width, distance and orientation angle of the grooves on the load-carrying capacity. Furthermore, this study also evaluates the effects of the bearing number and width-to-diameter ratio on the load performance of micro-grooved TPFPAJBs. Copyright \u00a9 2015 John Wiley & Sons, Ltd. Lubrication Science 2016; 28:207\u2013220 DOI: 10.1002/ls The calculation coordinate for a TPFPAJB is shown in Figure 1, where xOby is the calculation coordinate system of the bearing, z is the axial coordinate of the bearing, Ob is the bearing centre, Oj is the journal centre, R is the journal radius, Rpad is the bearing radius, D is the journal diameter, B is the bearing width, \u03c9 is the rotating speed, \u03b8 is the deviation angle, fx and fy are the gas film forces acting on the journal along the negative x and y direction respectively, \u03b1 is the arc bushing angle of each bearing pad, \u03b6 is the groove width angle, \u03d5 is the angle which starts to calculate from negative y direction, \u03b2 is the arc bushing angle between negative y direction and the leading edge of the first bearing pad, \u03c6 is the dimensionless circumferential coordinate calculated from the start of the deviation line OjOb and h is the gas film thickness", + " Gas film thickness of the bearing pad The geometry model of a micro-grooved bearing pad is shown in Figure 3, where xs is the coordinate of the leading edge of the micro-grooved bearing pad along the x direction, lx is the length of the microgrooved bearing pad along the circumferential direction, \u03b3 is the orientation angle of the grooves, wg is the groove width, sg is the groove distance, hg is the groove depth and lg is the distance between two adjacent grooves. The dimensionless variables are defined in the following form. H0 \u00bc h0=c \u00bc 1\u00fe \u03b5 cos \u03c6; Hg \u00bc hg=c; Wg \u00bc wg=R; Lg \u00bc lg=R; Sg \u00bc sg=R; L1 \u00bc lx=R; L2 \u00bc B=R; \u03c6s \u00bc xs=R; (1:8) where h0 is the gas film thickness of the bearing pad without parabolic grooves, \u03b5 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2j \u00fe y2j q =c is the eccentric ratio, and xj and yj are the coordinates of the journal centre Oj along the x and y directions respectively (Figure 1). As developed elsewhere,1 for 0\u00b0\u2264 \u03b3<90\u00b0, the dimensionless gas film thickness of a micro-grooved bearing pad can be written as Copyright \u00a9 2015 John Wiley & Sons, Ltd. Lubrication Science 2016; 28:207\u2013220 DOI: 10.1002/ls H \u03c6; \u03bb\u00f0 \u00de \u00bc H0 \u00fe Hg 4Hg W2 g tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 \u03bb cos \u03b3 n1Lg Sg 1 2 Wg 2 \u03bb < tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 and \u03bb1 < \u03bb < \u03bb2 H0 \u00fe Hg 4Hg W2 g \u03bb tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de \u00fe 1 2 L2 cos \u03b3 n1Lg 1 2 Wg 2 \u03bb > tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 and \u03bb3 < \u03bb < \u03bb4 H0; elsewhere 8>>>>>>>>>< >>>>>>>>>: (1:9) where n1 \u00bc fix tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2L2 \u03bb\u00bd cos \u03b3 Lg , fix is a function rounding a number to the nearest integer towards zero, \u03bb1 \u00bc tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 n1 \u00fe 1\u00f0 \u00deLg sec \u03b3, \u03bb2 \u00bc tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 n1Lg \u00fe Sg sec \u03b3, \u03bb3 \u00bc tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 \u00fe n1Lg sec \u03b3 and \u03bb4 \u00bc tan \u03b3 \u03c6 \u03c6s\u00f0 \u00de 1 2 L2 \u00fe n1Lg \u00feWg sec \u03b3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002527_amm.856.231-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002527_amm.856.231-Figure3-1.png", + "caption": "Figure 3: Light weight gear wheel with filigree structures, designed for LBM manufacture [9]", + "texts": [ + " \u201cCost-benefit powder losses\u201d are based on economic considerations. Even though fractions of the powder are still within specifications, an isolation of these particles may be technically feasible, but not economically. A processing for reuse is therefore not practical from a cost-benefit point of view. In order to investigate material efficiency and powder losses for LBM, three scenarios are analyzed on an empirical basis. LBM machines were chosen that are suitable for production applications and therefore represent industrial practice. The same part (see figure 3) was manufactured on all machines and material losses according to table 1were measured. The part can be built without supports and due to its geometry no powder is enclosed within the final part, so that part inclusions do not occur. The LBM machines build platform sizes range between 125 x 125 mm\u00b2 and 250 x 250 mm\u00b2. Laser powers of 200 W or 400 W are applied. The machines were cleaned with a cloth for which the weight difference was measured before and after powder absorption. Part removal and build chamber cleaning were realized in two phases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure8-1.png", + "caption": "Figure 8. The relationship of non-circular gear and shaper cutter at point P.", + "texts": [ + " At any meshing point P, shown in Figure 7, it is a common tangent point of the three surface instantaneously and keeps pure rolling contact, respectively, the three surface are the surface 1 of shape cutter, the surface 2 of non-circular gear and the surface 3 of curve-face gear. So, it can be guaranteed that the non-circular gear can mesh with the curve-face gear correctly at any moment. Point P is the contact point of the contact line L12 on non-circular gear and the contact line L13 on curve-face gear in the process of transmission, which makes the contact pattern restricted to local and the meshing process more smoothly. Derivation of the line contact of non-circular gear As shown in Figure 8, in the meshing process of noncircular gear and shaper cutter, the meshing equation (10) can be simplified as follows f \u00f0 1, s\u00de \u00bc nf !\u00f01\u00de vf !\u00f012\u00de \u00bc 0 \u00f012\u00de nf !\u00f01\u00de is the unit normal vector and it is equal to nf !\u00f02\u00de, vf !\u00f012\u00de is the relative velocity of non-circular gear and shaper cutter in coordinate system Of XfYfZf. PB ! is the absolute velocity vector at point P of non-circular gear, and it is perpendicular to the radius vector O2P ! . PA ! is the absolute velocity vector at point P of shaper cutter, and it is perpendicular to the radius vector O1P ! . The relative velocity vf !\u00f012\u00de at point P can be represented by the two absolute velocity vectors shown in Figure 8. So, the relative velocity of point P can be established as follows v! \u00f012\u00de f \u00bc v! \u00f01\u00de f v! \u00f02\u00de f \u00bc PA ! PB ! \u00f013\u00de According to the mathematical equation of surface 1 shown in Figure 3, the unit normal vector nf !\u00f01\u00de of at The University of Melbourne Libraries on June 5, 2016pic.sagepub.comDownloaded from shaper cutter can be represented as follows nf !\u00f01 \u00de \u00bc @ r1 ! @ s @ r1 ! @us @ r1 ! @ s @ r1 ! @us \u00bc cos\u00f0 os \u00fe s\u00de sin\u00f0 os \u00fe s\u00de 0 2 64 3 75 \u00f014\u00de And the transmission matrix Mf10 , Mf2 can be represented as follows Mf10 \u00bc cos sin 0 0 sin cos 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 \u00f015\u00de Mf2 \u00bc cos 1 sin 1 0 0 sin 1 cos 1 0 0 0 0 1 0 0 0 0 1 2 6664 3 7775 \u00f016\u00de Based on the principle of coordinate transformation and equation (14), the absolute velocity of shaper cutter can be represented as follows vf " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000665_j.procir.2014.05.032-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000665_j.procir.2014.05.032-Figure1-1.png", + "caption": "Figure 1 Description of a single robot component", + "texts": [ + " Basically, the following guiding principle applies: For industrial manipulators a fast and safe algorithm for collision detection and avoidance is more important than exact analytical results. Therefore it is feasible to approximate the actual robot geometry with simplified structures which completely enclose the robot. In the present case this is realized as follows: For each axis a finite line is spanned through the two adjacent joints and . Also, a radius is introduced which describes the maximum distance between the line and the surface of the related robot component (see Figure 1). The describing equation for the line can be found in (2). )( 1 nnnnn PPPs (2) The point Pn represents the origin of a joint and is related to the world coordinate system Oo. Therefore a coordinate transformation is required. This can be done using the DenavitHartenberg representation [21]. Once the robot components are described, following conclusion is allowed: a collision exists if and only if the minimal distance between robot lines and is smaller than the sum of the related radiuses and . For fast moving robots we have to consider the whole systems response time", + " Maximum velocity calculation Given a point at two timestamps and with as relating time difference, the velocity in-between and can be linearly approximated as seen in equation (4). T critncritn n 1,, || PP v (4) For a safe estimation of imminent collision only maximum velocities are taken into account. Considering a linear relationship between velocity and distance to the point of initial movements the point have to be both located at maximum distance to the origin joint and inside the introduced bounding volume. So in case of the robot description model presented in Figure 1, the critical point is calculable with equation (5). 21 1 1, |||| nn nn nncritn r PP PPPP (5) 3.2. Minimum distance calculation This section will address how to calculate the minimum distance (see equation (3)). The calculation procedure thereby depends on the relative position to each other. Therefore, configuration tests are needed which allow to select the proper procedure. Figure 3 shows the projection of three general configurations. The minimum distance in case (A) corresponds to the shortest distance between the lines, in case (B) it is found as the shortest distance between a point and a line and in case (C) the minimum distance is equal to the shortest distance between two points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003851_sensors47125.2020.9278934-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003851_sensors47125.2020.9278934-Figure1-1.png", + "caption": "Fig. 1. A disruptive approach to manufacture force sensors using the laserbased powder-bed-fusion process to realize application-specific adaption and mechanical connection. Top: The lower base body is additive manufactured using laser-based powder-bed-fusion. Interuption of the process, allows to insert a steel plate, which is the measuring element carrier. Subsequently, the process is continued and the upper base body is printed to achieve a firmly bonded material connection between the base body an the steel plate resulting in a deformation element. Middle: The manufactured deformation element is applied with strain gauges. Bottom: Manufactured force sensor in full bridge configuration. The strain gauges are sealed with silicone.", + "texts": [ + " Therefore, we present a disruptive approach to manufacture force sensors based on the LPBF process. The additive manufacturing process potentially enables application-specific 978-1-7281-6801-2/20/$31.00 \u00a92020 IEEE Authorized licensed use limited to: Carleton University. Downloaded on June 01,2021 at 06:17:21 UTC from IEEE Xplore. Restrictions apply. adaption in terms of size and mechanical connection. This approach consists of integrating a conventionally manufactured stainless steel plate, which serves as the measuring element carrier, in a base body produced by LPBF (Fig. 1, top and middle). Material connection is created by the LPBF process. Application of the strain gauges is done after manufacturing the deformation element. Main aspects of this manufacturing method are the control of the thermal processes during manufacturing to minimize the induced stress in the built deformation element, the process stability and the assurance of a sufficient and reproducible strain transmission at the LPBFmanufactured connection point. In first tests we produced a basic deformation element of a one-dimensional cantilever-based force sensor and applied strain gauges in the center of the steel plate", + " The deformation element of our prototypes is based on load cells for measuring tensile and compressive forces. In the first step, a base body (stainless steel 1.4404) is built up to a height of 8 mm with the LPBF process. Then, the steel plate (stainless steel 1.4310, 50 mm x 10 mm x 2 mm) is integrated during interruption of the LPBF process in the provided cut-outs (5.1 mm x 10.1 mm x 2 mm). Subsequently, the process is continued to achieve a firmly bonded material connection between the base body and the steel plate (Fig. 1, top). We vary the exposure area (Fig. 2) to achieve different connection points from base body to steel plate and in order to investigate the strain transmission depending on the exposure area. This way, three deformation elements are manufactured. These deformations elements are reworked by milling and M5 threaded holes are drilled for fastening to the measurement setup. A finite element analysis (COMSOL Multiphysics 5.5, Burlington, USA) is carried out to find a valid region for the strain gauges for a nominal load without exceeding the yield strength", + "5/350, HBM/HBK, Darmstadt, DE) with a gauge factor of 1.96\u00b1 1.5 % and a base resistance of 350 \u2126\u00b1 0.30 %. The mechanical connection of the strain gauges is done by using a hot-curing phenolic resin (P250, HBM/HBK, Darmstadt, DE), after preparing the application area with sandpaper and subsequent cleaning with acetone. The single strain gauges are connected to a full bridge circuit using enameled copper wire. For further protection against environmental influences, e.g. humidity or dirt, a silicone seal (SG250, HBM/HBK, Darmstadt, DE) is used (Fig. 1, bottom). The built-up prototype force sensors are characterized using a universal testing machine (model inspekt Table5, Hegewald & Peschke, Nossen, DE) with a 100 N reference force sensor (accuracy 0.02 %) regarding their hysteresis and linearity with respect to a best-fit line (Fig. 4). This best-fit line goes through the initial point, such that the maximum deviations Authorized licensed use limited to: Carleton University. Downloaded on June 01,2021 at 06:17:21 UTC from IEEE Xplore. Restrictions apply" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002246_978-3-319-15684-2-Figure4.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002246_978-3-319-15684-2-Figure4.1-1.png", + "caption": "Fig. 4.1 Kinematic model of a biped robot", + "texts": [ + "), Advances in Mechanical Engineering, Lecture Notes in Mechanical Engineering, DOI 10.1007/978-3-319-15684-2_4 23 walking robots lean only by one foot on the ground for an appreciable period of time [3]. For a biped robot, the support area is small. Because it is passively unstable and non-linear it is not easy to design a walking controller. The control system should provide processing of information about the area [7, 11], making decisions about the movement, and control over the implementation. Figure 4.1 shows the structure of the biped and coordinates used to describe the configuration of the system. In [4, 5] we have obtained the equations of spatial movement, investigated the possibility of automatic control for stable and constant biped walking, and defined the desired time and place for touchdown at the end of the step and at the beginning of the next one. 24 V. Tereshin and A. Borina This paper proposes a method of biped walking control in different modes: walking in the up or down direction, walking up stairs or down stairs, and rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003272_j.promfg.2020.05.027-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003272_j.promfg.2020.05.027-Figure7-1.png", + "caption": "Figure 7. Harmonic response analysis results at the resonant frequency with the displacement vectors.", + "texts": [ + " An exciting voltage of 45 Vrms (Root-Mean-Squared voltage), of about 64 Vp-p Peak-to-Peak voltage, was applied to the piezoelectric ceramic rings to simulate the actual power level in vibrations. As shown in Figure 6, the waveguide-based vibrator with the design parameters (geometric dimensions and waveguide paths) is able to deliver harmonic responses on both the longitudinal direction and the torsional direction, as shown in Figure 6. Figure 6 shows the vibrator has a resonant frequency at 25,514 Hz. Figure 7 shows a steady-state hybrid L&T vibration at the corresponding resonant frequency. The vibration displacement vectors represent the exciting energy propagating alone the six waveguides to the front end with directional changes. At the front end of the vibrator, a steadystate hybrid L&T vibration can be observed (see Figure 7). To validate the mechanical property of the vibrator, this paper also analyzes the maximum stress of the vibrator at the resonant frequency under different exciting voltages. The study would be important for dynamic applications, especially on fatigue behavior. In the FE analysis, the maximum principal stress method was used to analyze the stress distribution, as shown in Figure 8. The maximum stress increases as the applied driving voltage are increased, as shown in Figure 9. The maximum voltage is 60 Vrms based on the capacity of PZT stacks and the power" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure8.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure8.4-1.png", + "caption": "Fig. 8.4 Sketch of the virtual reference frames position at the contact points of each of the three omni-wheels of an omnidirectional robot", + "texts": [ + "7 Gravity effect over an free-floating rigid mass . . . . . . . . . . . . . . 262 Fig. 6.1 Linear and angular velocities associated to geometric origin g of frame R1 and linear and angular velocities of the center of mass cm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 278 Fig. 8.3 Left: Swedish wheel, also known asMecanum wheel, invented by the Swedish engineer Bengt Ilon. Right: Omni-wheel, an upgraded design of a omnidirectional wheel . . . . . . . . . . . . . . . 337 Fig. 8.4 Sketch of the virtual reference frames position at the contact points of each of the three omni-wheels of an omnidirectional robot . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 337 Fig. 8.5 Differential robot Pioneer 3DX from Adept Technology, Inc. for in-door purposes, [Photograph taken from http://www.mobilerobots.com/Libraries/Downloads/ Pioneer3DX-P3DX-RevA.sflb.ashx] . . . . . . . . . . . . . . . . . . . . . . 346 Fig. 8.6 Left: 3D sketch for the contact forces: vertical restrictive force, lateral friction and the power torque in the first wheel", + " Then, the gravity vector is g0 = \u239b \u239d 0 0 \u2212g \u239e \u23a0 (8.17) and the gravity wrench (\u2212MG(1)) is expressed by (6.103). 336 8 Model Reduction Under Motion Constraint The Exogenous Forces F Consider that this robot has three omnidirectional wheels, like in Fig. 8.3-right. For simplification purposes, consider the wheels to be mass-less bodies and the contact point between the floor and the wheel to be constant relative to the wheel position (refer to Fig. 8.2). For kinematic analysis, consider the sketch of Fig. 8.4, where the non-inertial frame is placed at the geometric center of the wheels such that the x-axis is positive to the front of the vehicle, the z-axis is positive in the upward direction and the y-axis is placed according to the right-hand rule. Consider three local frames ri at the contact point of each driving wheel i with the x-axis parallel to the rotation axis of each wheel, whose origin positions rri \u2208 R 3 are function of the radius R of the circle in which each wheel is placed equidistantly from each other, and the vertical position h of the local frame w.r.t the floor: 8.2 Model Reduction, the Dynamical Approach 337 338 8 Model Reduction Under Motion Constraint rr1 = \u239b \u239d R 2\u221a 3R 2\u2212h \u239e \u23a0 ; rr2 = \u239b \u239d \u2212R 0 \u2212h \u239e \u23a0 ; rr3 = \u239b \u239d R 2 \u2212 \u221a 3R 2\u2212h \u239e \u23a0 Also form the sketch in Fig. 8.4 the rotation matrices for each contact frame are defined as follows: Rr1 1 = Rz,\u03c0/3 = \u23a1 \u23a3 1/2 \u2212\u221a 3/2 0\u221a 3/2 1/2 0 0 0 1 \u23a4 \u23a6 ; Rr2 1 = Rz,\u03c0 = \u23a1 \u23a3 \u22121 0 0 0 \u22121 0 0 0 1 \u23a4 \u23a6 ; Rr3 1 = Rz,\u2212\u03c0/3 = \u23a1 \u23a3 1/2 \u221a 3/2 0 \u2212\u221a 3/2 1/2 0 0 0 1 \u23a4 \u23a6 The exogenous wrench F is the addition of the three wrenches produced by the 3 wheels: F = X T r1 F (r1) c1 + X T r2 F (r2) c2 + X T r3 F (r3) c3 (8.18) where Xri = Rri i T (rri ), R ri 1 are the extended rotations that maps the coordinates of the frame ri to the local frame 1, and T T (rri is the extended translation that displace wrenches from the contact point to the vehicle\u2019s frame", + "91) where \u03c4 r0 \u2208 R q represents some redundant restriction forces, in the same direction that \u03c4 r , that may exist in the system but have no influence in its dynamics. Remark Notice that the reduced model with the restrictive force vector can be computed as ( \u02d9\u0304\u03bd \u03c4 r ) = [ M\u0304\u22121ST 0 \u2212 +UT PS Q ]( E \u03c4\u0304 \u2212 h(\u03b8, S\u03bd\u0304) \u2212 MS\u0307\u03bd\u0304 \u03c4 r0 ) (8.92) 8.3.3 Example 3: The Omnidirectional Mobile Robot, Kinematic Approach Consider again theomnidirectional robotRobotinoXT, fromFesto, showed inFig. 8.1. The omnidirectional wheels position sketch is shown in Fig. 8.4 and the dynamic equation is given after expression (8.9) as: M \u03bd\u0307 \u2212 T (\u03bd)M\u03bd \u2212 mT T (rc)G(\u03b8) = E \u03c4\u0304 + G\u03c4 r where the reduced generalized force \u03c4\u0304 (i.e. the 3 torques provided at each wheel axis) and the restricted generalized force \u03c4 r (i.e. the 3 vertical contact forces at each 360 8 Model Reduction Under Motion Constraint wheel), are given by (8.19); and the operators E and G are given by expressions (8.20): \u03c4\u0304 = \u239b \u239d \u03c4m1 \u03c4m2 \u03c4m3 \u239e \u23a0 ; \u03c4 r = \u239b \u239d fz1 fz2 fz3 \u239e \u23a0 ; E = 1 r \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u221a 3/2 0 \u2212\u221a 3/2 1/2 \u22121 1/2 0 0 0 h/2 \u2212h h/2 \u2212\u221a 3 h/2 0 \u221a 3 h/2 \u2212R/2 R \u2212R/2 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 ; G = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 0 0 0 0 0 0 1 1 1\u221a 3R/2 0 \u2212\u221a 3R/2 \u2212R/2 R \u2212R/2 0 0 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 Definition of matrices S and U " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure16-1.png", + "caption": "Figure 16. Experimental setup to measure the static frictional resistance of each gear type, including that with flat passive rollers (The gear in this picture is a normal spur gear for comparison with the roller gear)", + "texts": [ + " This meant the advantage of the gear with conical passive rollers was also confirmed in terms of dynamic frictional resistance. The static and dynamic frictional resistances of the gear with flat passive rollers were also compared with other gear types in this study. Virtually the same procedure and experimental setups as for the experiments involving the gear with conical passive rollers in former chapters were used for those featuring the gear with flat passive rollers. The experimental setups to measure static frictional resistances are shown in Fig. 16. In these experiments, we used the real planar omnidirectional driving gear because the gear with flat passive rollers can slide continuously between the intermittent teeth of the omnidirectional driving gear. In the experiment to measure static frictional resistance, the weight on the pulley on the left of Fig. 16 increased continuously, while the translational force on the omnidirectional gear when the gear started to slide was measured by the force gauge on the left. The teeth width of the gears compared with the gear with passive rollers is 20 mm, and their module and the number of teeth are the same as the gear with passive rollers. The X-axis of the graph of Fig. 17 is the pressing force from the force gauge on the right. The average values after the experiments had been repeated 10 times were plotted on the graph, and the error bars show the standard deviation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003589_1350650120962973-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003589_1350650120962973-Figure1-1.png", + "caption": "Figure 1. Multidegree-of-freedom dynamic gear model of the global mechanical system.32", + "texts": [ + " The RMS of the power losses in dynamic stabilized conditions using a local coefficient of friction using a constant and variable surface roughness are compared with experimental results and with the one predicted using the classical methods (constant COF). The simulation results validate the higher accuracy of the new proposed model based on EHL formulation considering profile errors against other models. The proposed model is an extension of that described in32 to include the influence of frictional effects (average COF (constant) and local COF (time dependent l(t) (see equation (10))) and tooth shape deviations (e\u00f0Mi\u00de) via gears power loss. Figure 1 shows a multi-body lumped parameter dynamic model with twelve degrees of freedom used to describe the spur geared gearbox system. This dynamic model consists of four inertia elements, i.e. a spur gear pair including a pinion and a gear, a motor and a load. The gear system is accounted for 4 nodes and to each node is associated 3 degrees of freedom. Shafts are decomposed into classic two-node beam elements with 3 DOFs per node which account for 2 translational and 1 torsional displacements. Each of the pinion and gear is modelled by onenode element with 3 DOFs per node, which represent the infinitesimal translations and rotations superimposed on rigid cylinder motions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003181_tase.2020.2993277-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003181_tase.2020.2993277-Figure5-1.png", + "caption": "Fig. 5. Hardware system of the LDNSA prototype.", + "texts": [ + " It is necessary to modulate Kp and Kd to satisfy the abovementioned two situations according to further simplify the inequations. Therefore, this demonstrates that the state of the system is asymptotically stable when the impedance and feedback controller parameters selected are satisfied with (34). In this section, whether the hysteresis compensation algorithm contributes to overcome hysteresis is discussed. Thus, a series of visualized experiments is presented in the following. First, a hardware system is constructed, as shown in Fig. 5. DSP TMS320F28335 made by the TI Company is used as the core controller of the entire system. The deformation of the elastic component is measured by a magnetic linear encoder, which is an MSR5000 with 0.005-mm resolution. A ZNNT torque sensor (0\u2013100 Nm) is also used in the system to obtain the torque exerted on the actuator. The motor combination contains a MAXON EC motor (397172), gearbox (gear ration: 66:1), and motor-driven card (ESCON 50/5). Position signals of the MAXON EC motor encoder are acquired via the quadrature encoder pulse (QEP) module of the digital signal processor (DSP)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000898_10402004.2013.830799-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000898_10402004.2013.830799-Figure2-1.png", + "caption": "Fig. 2\u2014Schematic diagram of a cylindrical roller bearing.", + "texts": [ + " Having obtained the load distribution, the individual stiffness and damping coefficients of the inner and outer race-to-roller contact of the load sharing rollers are determined using the curvefitted relations developed in part I (Chippa and Sarangi (22)). The relations of stiffness and damping are K\u0304 = a1W\u0304k1 U\u0304k2 Gk3 C\u0304 = b1 r\u0304 W\u0304c1 U\u0304c2 Gc3 . [5] The constants and exponents of various dimensionless parameters in the above equation are given in Table 6 of part I (Chippa and Sarangi (22)). The stiffness and damping coefficients of the inner race\u2013 roller\u2013outer race contact are modeled as a series connection of two linear spring\u2013dampers as shown in Fig. 2. The equivalent stiffness and damping of a roller is determined by the following equation: Kr + j \u03c9bCr = ( 1 Ka + j \u03c9bCa + 1 Kb + j \u03c9bCb )\u22121 , [6] where \u03c9b is the average of inner and outer race ball pass frequencies. The overall stiffness (Kr) and damping (Cr) matrices of a roller bearing are assembled using compatibility and coordinate D ow nl oa de d by [ U N A M C iu da d U ni ve rs ita ri a] a t 1 8: 49 2 5 D ec em be r 20 14 1100 S. P. CHIPPA AND M. SARANGI transformation as follows: Kr = z\u2211 i=1 Kr i N T i Ni Cr = z\u2211 i=1 Cr i N T i Ni , [7] where z is the number of load sharing rollers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002991_ccwc47524.2020.9031137-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002991_ccwc47524.2020.9031137-Figure4-1.png", + "caption": "Figure 4. The Rplidar 3600 Laser Range Scanner [2]", + "texts": [ + " 0737 Authorized licensed use limited to: Fondren Library Rice University. Downloaded on May 18,2020 at 00:04:10 UTC from IEEE Xplore. Restrictions apply. The hardware platform shown in Fig. 3 [2] of the robot is a modular design, which is designed with six different modules like a mobile module, module for laser-scanning, somatosensory module, module for the arm, and modules for display and voice. This robot is build using cost-effective sensors and uses a controller such as the Rplidar Al shown in Fig. 4 [2] used for location and navigation, Kinect (used to get the gestures of human for controlling the robot arm) and Arduino board adapted for the Robot Operating System (ROS). a) Sensors b) ROS framework in the computer c) Actuator The sensors gather the state of the robotic system and other environmental information to the ROS based setup framework. The ROS is the communication channel of the entire system. It processes the data and makes the decision for the output devices such as actuators. The actuator performs the necessary actions and displays the information as needed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002902_j.cirpj.2020.01.001-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002902_j.cirpj.2020.01.001-Figure4-1.png", + "caption": "Fig. 4. Results of the mathematical modeling process in graphical form at various time p second layer; (d) during the third layer.", + "texts": [ + " Point 0 coincides with the point around which the beam oscillates. Point 10 is located at approximately 5,6 mm behind the beam and point 10 is located at 5,6 mm in front of the beam. Between the points 10 and 10 were also selected control points with the indices 8, 6, 4, 2, 0, 2, 4, 6, 8 (they are not shown in the Fig. 3) equidistant from each other. Please cite this article in press as: D.A. Gaponova, et al., Effect of reheatin wire deposition method, NULL (2020), https://doi.org/10.1016/j.cirpj.20 Fig. 4 shows the mathematical modeling results in a graphical form. It can be seen that a stationary regime is not observed because of the deterioration of heat removal from the product and change in massiveness of the body. The molten pool length\u2014represented by grayscale\u2014increases from 10.5 mm when the first layer is deposited, to 18 mm when the third is complete. Fig. 5, a shows the temperature-time dependencies for each of the points (T\u2013100, T\u201350, T\u201310, T-2, T0, T5, T10) in Fig. 3\u2014which were stored at each time step" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002938_1687814020908422-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002938_1687814020908422-Figure5-1.png", + "caption": "Figure 5. Gear equivalent contact model.", + "texts": [ + " h* denotes the equivalent viscosity of the fluid, which can be obtained based on Ree\u2013Eyring model28,29 1 h = 1 h t0 tm sinh tm t0 \u00f012\u00de where t0 and tm denote the lubricant\u2019s reference shear stress and the viscous shear stress, respectively. h refers to the viscosity of the fluid. Because the EHL model mainly focuses on the contact area of the gear pair, a simple equivalent model is proposed to describe the contact area of the gear pair before the calculation of the fluid thickness, as shown in Figure 5. In the figure, K is the meshing point of pinion and gear. B1B2 indicates the actual mesh length of the gear pair, and N1N2 refers to the theoretical mesh length. v1 and v2 represent the rolling speeds of the pinion and gear, respectively. FN denotes the meshing force and w represents the tooth width. Req and Eeq stand for the equivalent curvature radius and Young\u2019s modulus, which can be obtained as 1 Req = 1 N2K + 1 N1K 1 Eeq = 1 2 1 m2 1 E1 + 1 m2 2 E2 \u00f013\u00de where m1 and m2 are Poisson\u2019s ratio of the gear pair" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001001_j.protcy.2014.08.013-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001001_j.protcy.2014.08.013-Figure4-1.png", + "caption": "Fig. 4: Model of a hexapod going up the incline", + "texts": [ + " This may be due to actuator losing power, brakes applied by implemented failure detection software or the direct failure of the joint. Number of degrees of freedom reduces by one if a single joint fails. As the leg is considered to have three degrees of freedom in normal state, a locked joint failure results in two degrees of freedom i.e. two dimensional motion. This constrains the workspace of the leg and there is specific range of kinematic constraints which the configuration of the failed leg must satisfy to guarantee a fault -tolerant gait. [2, 3] When a robot body has to climb upon inclined plane as shown in fig.4, it will be susceptible to toppling and sliding and more instability as compared to plane terrain motion. In such a condition, it is needed to select a particular gait pattern which will be fast and will have good stability margin, s. Now, as discussed in [1, 2, 3], we have different gaits available to choose from. Tripod gait has duty factor of 1/2 and is the fastest amongst all gaits. But stability margin s < 0 [1] and also it is statically unstable and possible only when the body is moving. Quadruped gait has stability margin s \u2264 0 and hence we have to select either quadruped or pentaped gait" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002560_cec.2016.7748339-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002560_cec.2016.7748339-Figure5-1.png", + "caption": "Fig. 5 Schematic of a CSTR system [27]", + "texts": [], + "surrounding_texts": [ + "The nonlinear differential equations are stated as [24, 25]\n)()()( 21 ttxtx \u03c9+= (16)\nM )()( M h)( M K)( 2 d 1 )(0 2 1 tutxtxetx tx +\u2212\u2212= \u2212 (17)\n)()()( 2 ttxty \u03c5+= (18)\nThe process and measurement noises, ),0(~)( 2 \u03c9\u03c9 Nt and\n),0(~)( 2 \u03c5\u03c5 Nt , are white. The sampling time is 0.4 sec as in [24]. Other parameters of the system are set as m 2(0)x1 = , sm 7.1(0)x2 = , N 6.1u(0) \u2212= , 0.001=\u03c9 and\n0.8=\u03c5 . Parameters of CACC are listed in Table 1. These parameters are tuned in a trial and error procedure. Time response of the states and the control signal are depicted in Fig. 3. Fig. 4(a) illustrares the estimated and true position of the carriage. The estimated, true and measured output are represented in Fig. 4(b). Results show that CACC can properly estimate the states and control the position.\nHere, )(tqc represents the coolant stream, )(tC and )(tT are the current concentration and temperature, respectively. Also, fC and 0T represent the feed concentration and the feed temperature, respectively. Table 2 represents constant parameters of CSTR. The stochastic differential equations of this system are stated as follows [27]:\n( ) )()(k)(C V q)( )(R\nE 0f tetCtCtC tT \u03c9+\u2212\u2212= \u2212\n(19)\n2016 IEEE Congress on Evolutionary Computation (CEC) 5123", + "( )\n( ))(Te1)( VC\nC\ne)( C Hk)(T V q)(\nco )(C\nhA\np\npcc\n)(R E\np\n0 0\npcc tTtq\ntCtTtT\ntq c\ntT\nc \u2212\u2212+\n\u2212\u2212=\n\u2212\n\u2212\n(20)\n)()()( ttTty \u03c5+= (21)\nwhere )0,(~(t) 2 \u03c9\u03c9 N and )0,(~(t) 2 \u03c5\u03c5 N are the process and observation noises, respectively. They are zero mean white noises. The sampling time is 0.05 min as in [27], and other parameters of the system are set as 601 \u2212=\u03c9 ,\n0.5=\u03c5 , 0.06C(0) = mol/l , 449T(0) = K , and 120(0)qc = min/1 .\nTable 2 Constant parameters of CSTR problem [27]\nParameter Value Unit Description\npcp C,C 1 kcal/g Specific heat coefficients\n,c 1000 g/l Liquid densities fC 1 mol/l Feed concentration\nH 5102\u00d7\u2212 cal/mol Heat of reaction\nhA 5107\u00d7 kcal/min Heat transfer term\nq 100 1/min Process flow rate 0k 10107.2\u00d7 1/min Reaction rate constant 0T 350 K Feed temperature coT 350 K Inlet coolant temperature\nE/R 4101\u00d7 K Activation energy term\nV 100 l Reactor volume\nThe objective is to track a desired concentration. The desired concentration and the corresponding reactor temperature are 0.1 mol/l and K 438.5 , respectively. The performance of CACC is evaluated through numerical simulation. Fig. 6(a) compares the true and estimated concentration. Fig. 6(b) shows the control signal. The estimated and true temperature is comparied in Fig. 7(a). Fig. 7(b) compares the estimated and measured temperature.\n0 1 2 3 4 5 6 0.06\n0.07\n0.08\n0.09\n0.1\n0.11\n0.12\nTime (min)\nC (m\nol /l\n)\nEstimation True Signal\n(a)\n0 1 2 3 4 5 6 102\n104\n106\n108\n110\n112\n114\n116\n118\n120\nTime (min)\n(b)\n5124 2016 IEEE Congress on Evolutionary Computation (CEC)", + "C. Guidance Problem\nAnother problem to evaluate the proposed controller is a nonlinear guidance problem. Fig. 8 shows two-dimensional engagement of a pursuer and a moving target. The motion equations of this system are stated as follows [28], [29]:\n)()( 21 txtx = (22)\n)()()( 2 312 txtxtx = (23)\n)( )( )()( )()(2\n1\n4\n1\nT\n1\n32 3 tx tx tx a tx txtxx \u2212+\u2212= (24)\n)(1)(1)( 44 tutxtx \u03c4\u03c4 +\u2212= (25)\nwhere Rtx =)(1 is the relative distance, between the\npursuer and the target, Rtx =)(2 , \u03bb=)(3 tx is the line of sight (LOS) rate, P4 )( atx = represents the pursuer acceleration normal to the LOS, Ta is target acceleration normal to the LOS, \u03c4 is the time constant of the pursuer, and u is the input of the system (Guidance Command).\nThe available measurements are the relative distance R and the LOS rate \u03bb ; therefore,\n)()( 0100 0001 )( ttt xy += (26)\nwhere T 4321 ][)( xxxxt =x is the state vector, )(ty is the observation vector, and )0,(~(t) 2 \u03c5N is the vector of observation noise. The objective is to find the optimal guidance command )(tu to nullify \u03bb .\nThe target is flying with constant velocity of 480 m/s. The pursuer is initially 3000 m away from the target with an initial velocity 913 m/s. The sampling time and the time constant \u03c4 are 0.1 sec. Also, it is assumed that the target has no maneuver. Thus, Ta is zero. The parameters of the system, needed for simulation, are set as m 3000)0( =R ,\nsR m 600)0( \u2212= , rad 015.0)0( =\u03bb , 2 P sm 0)0( =a and\ns108)0( 2mu = . Also, the variance of the measurement noises are set as 302 =R\u03c5 and 62 10\u2212=\u03bb\u03c5 .\nThe engagement trajectory is depicted in Fig. 9. The resultant miss distance is 0.94 m. Fig. 10(a) compares true, estimated and measured LOS rate. Fig. 10(b) illustrates the LOS rate for various lengths of the prediction horizon\n)12 8, 3,T( p = . As shown in Fig. 10(b), the rise time is decreased for 3Tp = (solid line). Fig. 11 demonstrates the pursuer acceleration and the guidance command. Comparison of the estimated and true relative distance is depicted in Fig. 12(a). Fig. 12(b) compares the estimated and measured relative distance. Finally, the estimation of R is demonstrated in Fig. 13.\nThe results, obtained from the above three experiments, show that the proposed nonlinear state estimation and control algorithm can successfully estimate the states of these systems and control the desired state with acceptable tracking error, control effort and control smoothness.\nV. Conclusion\nIn this paper, a new heuristic sample based controller was proposed for nonlinear stochastic systems. The proposed controller, called CACC, models the state estimation and control problems as a dynamic optimization problem and an optimization scheme, based on ant colony optimization, was utilized to solve this problem. An augmented model was proposed to integrate the estimation and control problems. An integrated cost function was proposed to minimize simultaneously the state estimation error, tracking error, control effort and control smoothness. CACC was tested over three nonlinear problems: a nonlinear cart and spring system, a nonlinear continuous stirred tank reactor, and a nonlinear two dimensional guidance problem. The results demonstrate that the new approach can simultaneously estimate the states and the optimal controls of a nonlinear stochastic system.\n2016 IEEE Congress on Evolutionary Computation (CEC) 5125" + ] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.30-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.30-1.png", + "caption": "Figure 9.30 An RPRPR2 manipulator in its various views.", + "texts": [ + "1 = [s3w\u2032 1 + b2(w\u2032 2 \u2212 \ud835\udf0e \u2032\u2032 4 d5?\u0307?4)]\u2215s3 (9.421) ?\u0307?5 = \ud835\udf14\u2217 2 \u2212 \ud835\udf0e \u2032\u2032 4 ?\u0307?1 = (s3\ud835\udf14 \u2217 2 + d5?\u0307?4 \u2212 \ud835\udf0e\u2032\u20324 w\u2032 2)\u2215s3 (9.422) The above equations have the following implications in the task and joint spaces. (i) Task space motion restriction implied by Eq. (9.417): \ud835\udf14\u2217 3 = 0 (9.423) (ii) Joint space motion freedom implied by Eq. (9.419): ?\u0307?4 + ?\u0307?6 = \ud835\udf14\u2217 1 (9.424) Owing to this freedom, one of ?\u0307?4 and ?\u0307?6 can be selected arbitrarily. Then, Eqs. (9.420)\u2013(9.422) give ?\u0307?1, s\u03073, and ?\u0307?5 accordingly. Figure 9.30 shows an RPRPR2 manipulator in its various views. As noticed, its wrist is not spherical, because the fourth joint is prismatic. It is another conceptual manipulator like the one studied in Section 9.5. It is studied here as a typical example of a manipulator that looks as if it does not have an analytical inverse kinematic solution. Yet, it is shown that the solution can still be obtained analytically in an unusual way, which is similar to the way of obtaining the first-order semi-analytical solution for the manipulator of Section 9.5. (a) Rotation Angles \ud835\udf031, \ud835\udf032 = \ud835\udeff2 = 0, \ud835\udf033, \ud835\udf034 = \ud835\udeff4 = \u2212\ud835\udf0b\u22152, \ud835\udf035, \ud835\udf036 Four of the rotation angles (\ud835\udf031, \ud835\udf033, \ud835\udf035, \ud835\udf036) are joint variables. The second rotation angle is associated with the cylindrical arrangement of the joints 2 and 3. So, it is taken to be zero as discussed in Chapter 7. Thus, u\u20d7(2) 1 is arranged to be parallel to u\u20d7(1) 1 . On the other hand, the fourth rotation angle is associated with the prismatic joint 4. As seen in Figure 9.30, the unit vectors u\u20d7(3) 1 and u\u20d7(4) 1 imply that \ud835\udeff4 = \u2212\ud835\udf0b/2 about u\u20d7(4) 3 . (b) Twist Angles \ud835\udefd1 = 0, \ud835\udefd2 = \u2212\ud835\udf0b\u22152, \ud835\udefd3 = 0, \ud835\udefd4 = \u2212\ud835\udf0b\u22152, \ud835\udefd5 = \u2212\ud835\udf0b\u22152, \ud835\udefd6 = \ud835\udf0b\u22152 (c) Offsets d1 = 0, s2 = OQ, d3 = 0, s4 = QR, d5 = 0, d6 = RP Four of the offsets (d1, d3, d5, d6) are constant. Two of the offsets associated with the prismatic joints (s2 and s4) are joint variables. (d) Effective Link Lengths b0 = 0, b1 = 0, b2 = 0, b3 = 0, b4 = 0, b5 = 0 (e) Link Frame Origins O0 = O,O1 = O,O2 = Q,O3 = Q,O4 = R,O5 = R,O6 = P (a) Link-to-Link Orientation Matrices C\u0302(0,1) = eu\u03031\ud835\udefd1 eu\u03033\ud835\udf031 = eu\u03033\ud835\udf031 (9", + " On the other hand, \ud835\udf0e\u20322 is known owing to the specification of \ud835\udf192 according to Eq. (9.485). Therefore, Eq. (9.488) implies that \ud835\udf0e\u20325 = \ud835\udf0e\u20322sgn(r2s\ud835\udf191 + r1c\ud835\udf191) (9.496) Equations (9.466) and (9.467) imply that a position singularity may occur if q12 = 0, i.e. if r1c33 \u2212 r3c13 = r2c33 \u2212 r3c23 = 0 (9.497) In the general solution described above, c33 \u2260 0. In that case, Eq. (9.497) implies further that r1 = r3c13\u2215c33 and r2 = r3c23\u2215c33 (9.498) However, this singularity is not likely to occur. This is because, referring to Figure 9.30, Eq. Pair (9.498) implies that the points Q and O get coincident, but this coincidence is not possible due to the physical shape of the second joint. On the other hand, as also described above, the special solution with c33 = 0 (i.e. with c\ud835\udf035 = 0) can be obtained without encountering any singularity. (a) Angular Velocity of the End-Effector with Respect to the Base Frame Let \ud835\udf14 = \ud835\udf146. Then, Eq. (9.436) leads to the following expression. \ud835\udf14 = ?\u0307?1u3 + ?\u0307?3eu\u03033\ud835\udf031 u2 + ?\u0307?5eu\u03033\ud835\udf031 eu\u03032\ud835\udf033 u1 \u2212 ?\u0307?6eu\u03033\ud835\udf031 eu\u03032\ud835\udf033 eu\u03031\ud835\udf035 u3 (9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000512_j.renene.2015.08.077-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000512_j.renene.2015.08.077-Figure2-1.png", + "caption": "Fig. 2. Scheme of main sensor configurations.", + "texts": [ + " The detailed operations of each step are explained in the right of Fig. 1. 3. Hardware configuration As the aforementioned, the measurement system consists of four data acquisition units located in tower bottom, nacelle, hub and met-mast respectively. The metrological parameters are captured mainly from the meteorological parameter sensors in the met-mast unit. Load and operation quantities are collected from the strain gauges and other analogue/digital sensors installed in the tested WT. Their configurations are shown in Fig. 2. The anemometer and other sensors for measuring meteorological parameters are located on a meteorological tower at the turbine's hub height. The meteorological tower is away 2.38 times of rotor diameter from the tested WT. Most of operational datasets are read out from the electric control box located in the tower base. The rotor position is measured with an inductive sensor located on the rotor flange. The absolute position of the rotor is determined by combining this signal with the digital pulses from the rotating angle sensor. The strain gauges for measuring blade bending loads are mountedwithin the blades to avoid the harms from lightning or other harsh environments, which are located at 1e180 (flapwise) and 90 e270 (lead lag) directions of the same plane, 1.2 m away from the flange, as shown in Fig. 2. Since the local strain in connection gap between blades is very unstable, the measuring direction of the blade flap-wise bending is deflected 15 from the 90-degree direction for minimizing the effect of cross sensitivity. The strain gauges formeasuring rotor tilt moment and yaw bending are installed behind the spindle flange. The rotor bending measuring directions are similar to those of blade bending. The torque sensor for measuring rotor torque is just next to the strain gauge at the 0-degree direction, nearly being consistent with the 0- degree direction of rotor azimuth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001084_s12206-015-0217-8-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001084_s12206-015-0217-8-Figure4-1.png", + "caption": "Fig. 4. An example of stress extraction lines for the sensitivity analysis surface lines otal stress distributions on the inner surface lines.", + "texts": [ + " Total stress distributions on the inner surface lines have fluctuation characteristics and locations of the maximum total stress generation Table 1. Geometric variables of the BMI penetration no-zzles of all the domestic Westinghouse domestic PWRs in South Korea. Fig. 2. Basic dimensions of the geometric variables of the BMI penetration nozzle. points depend on the FE analysis variables. So, the sensitivity analysis used the total stress distributions extracted from the top point S along the inner surface lines, as shown in Fig. 4. Weld zone means an inner surface region of the nozzle corresponding to heights of the J-groove weld. Weld root means a point on the no-zzle inner surface corresponding to the lowest point of the J-groove weld. Fig. 5 depicts the total stress distributions on the inner surface lines (fn = 23.1o, tn = 13.38 mm, ro = 19.57 mm, qw = 25o, dw-up = 18.613 mm, dw-down = 31.7 mm, ww = 8.82 mm). The total stresses mean operating stresses at SS during NOP including welding residual stresses. As depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure21-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure21-1.png", + "caption": "FIG. 21. Deformed cross-sections for the case of N > 2, showing the natural shape.", + "texts": [ + " The experimental displacement, \u03b4\u2217 1 , is measured from the supports to the contact point of the indenter and the creases. Using the geometry of this configuration, the relationship between experimental and model displacements found: \u03b4\u2217 2 = \u03b4 \u2212 di 2 ( \u03b4 ) = \u03b4 ( 1 \u2212 di 2 ) . (B2) When the number of creases is greater than two, the creases do not initially lie in the same plane as the supports. As a result, the experimental displacement \u03b4\u2217 3 is measured from the shape in the lowest energy stable state, as shown in Fig. 21. The model displacement, \u03b4, is measured from the crease edge to the central vertex. When the deformation of the disk raises the central vertex above the supports (\u03b4 < 0), the contact between the indenter and the disk is at the hole edge, therefore \u03b4\u2217 3 = ( L 2 tan \u2212 r sin ) \u2212 ( L 2 tan 0 \u2212 r sin 0 ) , where the first bracketed term is the distance from the support to the edge of the hole in the deformed state and the second bracketed term is the distance from the support to the hole edge in the natural state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002251_b978-0-12-803581-8.03009-5-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002251_b978-0-12-803581-8.03009-5-Figure1-1.png", + "caption": "Figure 1 The formation of plastic hinges in an open-cell foam.", + "texts": [ + " The simplest analysis involving plasticity could be developed for the collapse of cell walls leading to the estimation of maximum allowable remote stress that a piece of cellular material can withstand. This is taken up first in Section 1. Both open-cell and close-cell foam properties are reviewed. A brief summary of plastic response and recoil calculations for two-dimensional honeycomb structures is presented in Section 2. Plastic collapse in an open-cell foam occurs when the bending moment exerted on the cell walls exceeds the fully plastic moment (Thornton and Magee, 1975; Gibson and Ashby, 1997), creating plastic hinges like those shown in Figure 1. Closed-cell foams are \u2606Change History: July 2015. A. Bonfanti and A. Bhaskar updated the article by implementing Figures 5 and 6, along with the Plastic response and recoil of cellular materials section and rewriting the opening paragraph with new material. Reference Module in Materials Science and Materials Engineering doi:10.1016/B978-0-12-803581-8.03009-5 1 more complicated; in them, the plastic-collapse load may be affected by the stretching of cell faces as well as by the bending of the cell edges, and by the presence of a fluid within the cells. The simpler case of open cells is examined first. Plastic collapse occurs when the moment exerted by the force F (Figure 1) exceeds the fully plastic moment of the cell edges. For a beam with a square section of side t, this fully plastic moment is: Mp \u00bc 1 4 syst3 \u00bd1 where sys is the yield strength of the cell wall material. If the force F has a component normal to the cell edge (of length l), the maximum bending moment is proportional to Fl. The stress on the foam is proportional to F/l2. Combining these results gives the plastic-collapse strength of the foam: s plp Mp l3 For open-cell foams, the relative density r /rsp(t/l)2 (where r is the density of the foam and rs is that of the solid of which it is made) giving the collapse stress of the foam s pl: s pl sys \u00bc C1 r rs 3=2 \u00f0open cells\u00de \u00bd2 where the constant C1 contains all the constants of proportionality" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001071_s11740-013-0466-2-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001071_s11740-013-0466-2-Figure4-1.png", + "caption": "Fig. 4 Mathematical model of the grinding wheel", + "texts": [ + " The requirement for the analysis of the influence of kinematic modulated grinding processes on the surface structure is an initial provision of a mathematical modeling for 3D surface structures by using an overlapping movement. The approach presented in this work is based on a kinematic-geometrical model to realize a kinematic modulation, which essentially takes the approaches of modeling from Werner [11] and Zitt [12] into account. The basis for the mathematically modeling is a discretization of the grinding wheel and the workpiece in segments and points. The grinding wheel is assembled in n segments with a width bs. Figure 4 depicts the rough discretization of the grinding wheel and the workpiece. A stochastic distribution and formation of octahedron-shaped grains are used on the surface of all grinding wheel segments in which cutting edge depth z, grain distance lg and the grain orientation are varying stochastically. In addition, the grit concentration C by Zitt [12] is also taken into account. Grinding wheel wear, elastic and plastic deformation in chip formation as well as machine stiffness are not considered at first" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000754_cgncc.2014.7007393-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000754_cgncc.2014.7007393-Figure5-1.png", + "caption": "Figure 5. UAV landing virables schematic diagram", + "texts": [ + " The system compares the aircraft states with the statistical distribution of the real landing states, then calculate the probability of success landing, and make a missed approach decision. The flight control system is designed to track the aircraft aft, when the aircraft fly over the aft, it should maintain proper clearance and impact velocity, and arrest the cables success finally. Therefore the main wave-off decision parameters are Rh\u0394 , TDh\u0394 and IV\u0394 . The terminal control problem is shown in Figure 5. The UAV impact velocity is IV , the UAV height over the ramp is Rh , the UAV distance apart from the desired touchdown point is TDX . ( )IV\u03c3 \u0394 , ( )Rh\u03c3 \u0394 and ( )TDX\u03c3 \u0394 are the RMS distribution of the variable aforementioned. Distribution of these variables are Gaussian or standard distribution function, the real distribution measurements support the hypothesis of a normal distribution as showed in [11]. The probability of ramp-strike is 2 ( ) /2( ) ( )11 1 ( )2 R M R R h z R Mh h R hP e dz F h \u03c3 \u03c3\u03c0 \u2212\u0394 \u2212\u221e = \u2212 = \u2212 \u0394 , the probability of hard landing is ( ) 1 ( )R I M h I V P F V\u03c3 = \u2212 , and the probability of bolt is ( ) 1 ( )R TD M h TD I X P F X\u03c3 = \u2212 ,the probability calculation of the three variables is independent and mutually exclusive", + " It is noted that, in order to show the simulation result clearly, the total landing time after \u201ctip over\u201d in simulation system is longer than reality. The simulation is set in the last stages of landing, the aircraft\u2019s initial state was heading east, at a height of 100 meters, while the glideslope angle is 3 \u00b0. Deck motion compensation and noise rejection is applied in the flight control algorithm. The simulation results are shown in Figures 7 and 8. Figure 7 shows the state height h, speed V, angle of attack and angle of sideslip , while figure 5 shows the euler angle , , and their angular rate p, q and r. From the simulation results, we can see that in order to maintain the desired glideslope angle, initial height has been adjusted. Due to the initial setting speed deviation, aircraft fixes it quickly at the beginning. In the subsequent flight, the aircraft land successfully in stable attack angle and steady speed. The transient pitch angle of the final phase in figure 8 is that the aircraft adjusts its attitude on the deck. According to the simulation results, the system gets the expected results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002593_978-3-319-46155-7_6-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002593_978-3-319-46155-7_6-Figure5-1.png", + "caption": "Fig. 5 SLM machine concept of parallelization", + "texts": [ + " These investigations show that it is possible to heighten the theoretical build-up rate and the productivity significantly by using increased laser power up to PL\u00bc 2KW. Experts of the Cluster of Excellence designed and implemented a multi-scannerSLMmachine. Two lasers and two scanners are integrated in this system. These two scanners can be positioned to each other, either that both scan two fields on its own (double-sized build space) or that one scan field is processed with two scanners simultaneously (see Fig. 5). New scanning strategies can be developed and implemented with these multi- scanner systems (see Fig. 6). Scanning strategy 1: both scan fields are positioned next to each other with a slight overlap. This results in a doubling of the build area. By using two laser beam sources and two laser-scanning systems, both scan fields can be processed at the same time. In this case, the build-up rate is doubled. Scanning strategy 2: the two laser beam sources and the two laser-scanning systems expose the same build area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003680_icma49215.2020.9233573-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003680_icma49215.2020.9233573-Figure2-1.png", + "caption": "Fig. 2. the actual picture of the snake-like manipulator.", + "texts": [ + " Downloaded on May 30,2021 at 13:52:16 UTC from IEEE Xplore. Restrictions apply. The snake-like manipulator is driven by the rope, and the driving motor is separated from the manipulator body. The manipulator body is modularly combined. Each link is connected by a universal joint. We plan to build a snakelike manipulator with 16 degrees of freedom. The current laboratory prototype has 12 degrees of freedom, as shown in Fig. 1. Using the modified D-H method, the coordinate system corresponding to each joint is established, as shown in Fig. 2, and the D-H parameters of the snake-like manipulator are calculated to complete the forward kinematics solution of the manipulator. Obstacles are transformed into cylinders or spheres. Transforming obstacles into regular geometry in space, although the space of obstacles is expanded, the calculation is simplified, the calculation efficiency is improved, and the safety of the snake-like manipulator is guaranteed. The RRT algorithm needs to perform collision detection every time when it generates additional node", + " This paper presents a method for calculating the angle vector of the middle point. Firstly, the trajectory of the snake manipulator is generated from the initial point to 493 Authorized licensed use limited to: Carleton University. Downloaded on May 30,2021 at 13:52:16 UTC from IEEE Xplore. Restrictions apply. the expected point, and the generated trajectory is collision detected. The collision tracks are screened out from the generated tracks, and each obstacle corresponds to a collision track. By controlling the even joint angle change in Figure 2, the selected trajectory is changed. Judge the relative position of the obstacle center and the initial position of the mechanical arm. If the initial position is in the XY plane where the center of the obstacle is located, the even joint angle is reduced by 1 degree in each iteration until the generated trajectory does not collide, otherwise, it is increased by 1 degree. Finally, the collision free trajectory is generated as the middle point of the improved RRT algorithm. For multiple obstacles, multiple intermediate points are generated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000846_j.ymssp.2014.11.015-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000846_j.ymssp.2014.11.015-Figure6-1.png", + "caption": "Fig. 6. Impulse experiment to determine the viscous damping ratio associated with the lubricated contact regime. (a) Experimental setup; (b) top view of the dowel pin arrangement showing the three point contacts. Key: , contact point.", + "texts": [ + " Here, ve\u00f0t\u00de \u00bc _\u03c8 j\u00f0t\u00de\u00fe r \u0394\u00f0t\u00de sin \u03c6\u00f0t\u00de\u00fe\u03b1\u00f0t\u00de \u00fee sin \u03b1\u00f0t\u00de\u00fe\u0398\u00f0t\u00de _\u03b1\u00f0t\u00de\u00fe _\u0398\u00f0t\u00de : \u00f026\u00de The slide-to-roll ratio vr\u00f0t\u00de = ve\u00f0t\u00de \u00de could be easily changed by altering the geometry, such as PE ! , a, b and e. The modal damping ratio under lubrication depends on the oil viscosity and the materials in contact; hence it is determined experimentally using the half-power bandwidth method with both lubricants. The experimental setup consists of two masses (m1\u00bc1.4 kg and m2\u00bc1.8 kg) connected by three identical point contacts which are lubricated as shown in Fig. 6. These point contacts are obtained by placing three dowel pins (rd\u00bc3.2 mm) attached to m1 in one direction and two more dowel pins attached tom2 in the orthogonal direction, as shown. The system is placed on a compliant base (foam), and two accelerometers are attached to each mass. An impulse excitation is imparted to the system in the vertical direction with an impact hammer. The response accelerance spectrum of each mass along the vertical direction is then found by averaging signals from two accelerometers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure6-1.png", + "caption": "Fig. 6. Schematic view of lead-lag motion.", + "texts": [ + " Recently, we have been working on solving over one giga-degrees of freedom (DOFs) FSI problem using K computer [46]. In the future work, we will extend the present analysis to more realistic one which has a complex geometry. Although the extension engenders the increase of DOFs, it is straightforward to apply the analysis system. A flapping-wing motion can be described as a combination of three rotational degrees of freedom, namely stroke (yaw) (Fig. 4), pitch (feathering) (Fig. 5), and lead-lag (roll) (Fig. 6). According to one study [1], pitch plays a large part in the enhancement of lift generation. The pitch can be actively generated (active pitch) using actuators or passively generated (passive pitch) by the stroke if a deformable wing is adopted, as shown in Fig. 7. From a viewpoint of simplifying MAVs design, the passive pitch is very advantageous because of keeping the number of actuators low. The passive pitch has been investigated using experiments and numerical studies [15,17]. In the present study, we compare active and passive pitch motions quantitatively", + " The passive pitch angle is defined as the angle between two lines, namely a line on the rigid part of the wing chord and a normal line parallel to the y-axis. This angle is used to evaluate the degree of deformation, which depends on the flexibility of the upper part of the deformable wing. A lead-lag motion is observed, in terms of the wing tip trajectory, in several species of insects and vertebrates. The axis of lead-lag is the edge including the fixed point (0,0,0), which corresponds to the edge AB in Fig. 8b. Fig. 6 shows the lead-lag angle \ud835\udf03l defined by the following equation: \ud835\udf03l = CA(\u0398l sin(2\ud835\udf0bflt + \ud835\udefe) + D) (deg), (25) where \u0398l is the maximum lead-lag angle, fl is the lead-lag frequency, \ud835\udefe is the phase difference between stroke and lead-lag motions, and D is a constant for vertical translations. As mentioned in Section 2.2, in order to make our FSI analysis system robust to a wide range of flapping motions, we employ a specialized mesh under the SEMMT framework. In the SEMMT framework, if largesized elements are used for the outside of the SEMMT-applied part, we can keep the aspect ratio to some degree because of the following reasons" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure6.8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure6.8-1.png", + "caption": "Fig. 6.8 Vibration modes 2, 3, 4 and 5, taking the connection to the foundation into account", + "texts": [ + " Larger machines for higher speeds do have larger airgaps, machines with six and more poles have smaller air-gaps. Normal asynchronous machines should not have eccentricities of much more than 0.3mm. The eccentricity according fig. 6.7 causes a mode 7 and a mode 9. Mode 7 will have a frequency of twice the stator voltage supply frequency and the rotor slot modulated frequency. The amplitude is relative small in the range of 5000 N/m2. The radial forces cause vibrations of the stator or rotor core. The vibration will be a mixture of mechanical Eigen-modes. An overview of mechanical Eigen-modes is given in fig. 6.8. The stator gets a certain form, which defines the mode acc. literature [6.3, 6.6]. Often only free symmetric stators are shown. Indeed the mechanical design becomes asymmetric due to the connection of the motor to the foundation. One typical mode is split f.i in three. Only one of the multiple modes is shown in fig. 6.8 for each of the typical forms. 96 6 Noise Based on Electromagnetic Sources in Case of Converter Operation The calculation of the individual mechanical vibration amplitudes consists on the one hand of the determination of the Eigen-frequencies for the modes r of the mechanical structure and on the other hand on the determination of their amplitudes. The structure mechanical calculation has been described in literature [6.1, 6.3, 6.9]. Different approaches can be found in order to determine Eigen-frequencies and modes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002593_978-3-319-46155-7_6-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002593_978-3-319-46155-7_6-Figure9-1.png", + "caption": "Fig. 9 Mesh structure of optimization result including stress distribution", + "texts": [ + " Therefore, the dynamic load case is simplified to five static load cases, which consider the maximum forces at different times of the dynamic seat movement. Material input for the optimization is based on an aluminum alloy (7075) which is commonly used in aerospace industry: material density: 2810 kg/m3, E Modulus: 70.000 MPa, Yield Strength: 410 MPa, Ultimate Tensile Strength: 583 MPa and Poisson\u2019s Ratio: 0.33. The objective criterion of the optimization is a volume fraction of 15% of the design space. The part is optimized regarding stiffness. Figure 9 shows the optimization result as a mesh structure and an FEM analyzes for verification of the structure. The maximum stress is approx. 300 MPa, which is below the limit of Yield Strength of 410 MPa. Before manufacturing, the surfaces of the optimized part are smoothened to improve the optical appearance of the part. Compared to a conventional part (90 g), a weight reduction of approx. 15% (final weight 77 g, see Fig. 10) was achieved. Further improvements to increase the productivity of the process are needed for series part production" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002290_speedam.2016.7525820-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002290_speedam.2016.7525820-Figure5-1.png", + "caption": "Fig. 5. Structure of model B.", + "texts": [], + "surrounding_texts": [ + "Figs. 5 and 6 show the structure of model B and its cross section, respectively. The magnetized direction is indicated by the red arrows. In model B, ferrite magnets magnetized in the radial direction are inserted in the circumference of the shaft. As shown in Fig. 6, the ferrite magnets inserted in the upper core, and those inserted in the lower core are magnetized in opposing directions. In addition, the shaft is made of magnetic material, thereby acting as a three-dimensional magnetic path. As a result, the effective magnetic flux can be increased and the torque is also increased. Table V shows the total volume of ferrite magnet in model A (a/a\u2019 = 1/0.85) and model B. (a) Upper core (b) Lower core Fig. 6. Cross section of model B. N N N N S S S S Shaft (Magnetic material) N N N N N N S S S S S S S S S S N N N N Upper side of rotor core Lower side of rotor core Fig. 7. Structure of rotor. Fig. 8. Model C." + ] + }, + { + "image_filename": "designv11_34_0003847_0954405420978120-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003847_0954405420978120-Figure1-1.png", + "caption": "Figure 1. CAD model of a SCARA robot.", + "texts": [ + " Section 4 presents several simulation experiments of the proposed method and comparisons with other methods. Section 5 is the actual experiments by drawing with a TURIN SCARA robot. At last, Section 6 concludes this research. SCARA robot path local arc transition The representation of SCARA linear robot path In Cartesian space, a target point of linear robot path represents the position and orientation of Tool Coordinate System (TCS) relative to Robot base Coordinate System (RCS), as shown in Figure 1. The target points of SCARA robot path are represented by 4D point fPi xi, yi, zi, ui\u00f0 \u00degNi=0. Positions of SCARA robot can be represented by the Tool Center Point (TCP) (xi, yi, zi) in RCS. The rotation axis of the three rotation joints are constant, only rotation angles (ui) around z-axis of TCS are variables. The distance between two 4D points can be defined by position distance and orientation distance. Suppose a 4D point is Pi = xi, yi, zi, ui\u00f0 \u00de. Denote Pi, 3\u00bd = xi, yi, zi\u00f0 \u00de, Pi, u\u00bd = ui" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003203_j.est.2020.101417-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003203_j.est.2020.101417-Figure2-1.png", + "caption": "Fig. 2. Battery topologies for the 6S6P configuration.", + "texts": [ + " In this study, a battery having 373 Wh power and ~17 Ah capacity is to be designed by employing cylindrical cells with 2,8 Ah capacity. To achieve 17 Ah, six parallel branches are placed along with each having Table 1 Topology comparisons. Topology Mass (kg) Energy Density (Wh/ kg) Volume (m3) First Nat. Freq. (Hz) 6 \u00d7 6 3.48 107.18 2.567 10\u22123 987.7 4 \u00d7 9 3.24 115.12 2.242 10\u22123 2385.3 Circular 3.66 101.91 2.851 10\u22123 3212.1 3 \u00d7 12 3.24 115.12 2.261 10\u22123 3447.9 six series cells. The 6S6P battery is designed in four alternative topologies namely, 6 \u00d7 6, 4 \u00d7 9, Circular and 3 \u00d7 12 as illustrated in Fig. 2. Electrically identical, each configuration is compared with the others according to mass, volume, and the first resonance frequency criteria. Moreover, harmonic and random vibration simulation performances are assessed. Table 1 summarizes the mechanical and electrical performance criteria of the four topologies. It is clearly seen that the 3 \u00d7 12 topology alternative is the best option with one of the lowest mass (3.24 kg), the highest energy density (115.12 Wh/kg) and the highest natural frequency (3447 Hz)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003349_tmag.2020.3011612-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003349_tmag.2020.3011612-Figure4-1.png", + "caption": "Fig. 4. Cross section of the 48 stator slots, 4 poles DW and 30 rotor bars.", + "texts": [ + " ( )d L I t U t R I t e t dt (12) Where R and L are the resistance and inductance matrices given by (13) and (14) respectively. 3 0 0 0 0 0 0 bar bar bar end end end n R R R R R R R (13) , , , , 1 , 1, 3 0 0 0 0 0 0 0 0 0 0 i j i n i i j n n i n n n n n L M M M M L M M L (14) To validate the proposed approach (the equivalent electrical circuit and the calculation of its components), two different squirrel cage induction machines are studied. The first one is a distributed winding (DW) with 48 stator slots and 30 rotor bars as shown in Fig. 4. The second one is a fractional slot concentrated winding (FSCW) with 12 stator slots and 16 rotor bars as shown in Fig. 5. As the second winding is a double layer one, each stator slot houses two coil sets of either the same phase or two different phases. For each machines results are given for two operating points i.e. at synchronous speed (no load) and at load (slip s = 5%) and compared to the ones obtained by a 2D FEM in the same conditions while considering a linear B(H) curve of the magnetic material. A. Results for the DW machine Fig. 6.a shows the variation of the self-inductance of the first bar, L1, and Fig. 6.b the mutual inductances M1,2 and M1,10 versus the rotor position obtained with (10) and (11). Both present very low magnitude variations due to low stator slot opening, as shown in Fig.4, and hence low reluctance variation. Therefore, in a first assumption, they can be considered as constant whatever the rotor position, i.e. equal to their average values. This simplifies the matrix system (12) into: ( )dI t U t R I t L e t dt (15) Authorized licensed use limited to: East Carolina University. Downloaded on July 28,2020 at 15:04:57 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000185_ijpt.2018.090374-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000185_ijpt.2018.090374-Figure2-1.png", + "caption": "Figure 2 Schematic representation of gear teeth modification (a) Crowning (b) Tip relief", + "texts": [ + "23 y c e e e x R h U G W R \u2217 \u2212 \u23a7 \u23ab\u23a1 \u23a4\u23aa \u23aa\u239b \u239e= \u2212 \u2212\u23a2 \u23a5\u23a8 \u23ac\u239c \u239f\u23a2 \u23a5\u239d \u23a0\u23aa \u23aa\u23a3 \u23a6\u23a9 \u23ad (10) where the dimensionless groups are expressed as: ( )0 0 02 2, , , 4 2e e e r c r x r x x \u03c0\u03b7 U \u03c0W hU W G E h E R E R \u03c0 R \u2217= = = =\u03b1 Therefore, the generated viscous friction, using equation (9) becomes: vf \u03bcW= (11) The planetary wheel hub gears of the JCB Max-Trac rear differential is studied here. The input torque to the sun gear from the differential is 609 Nm at the speed of 906 rpm. Firstly, the effect of gear teeth longitudinal crowning is considered. Secondly, the influence of gear teeth tip relief modification is considered. Two parameters are involved in the tip relief modification: the extent of tip relief and the length of the relieved region. The results are presented for the cyclic meshing power loss of both the axle\u2019s planetary wheel hub sets. Figure 2 is a schematic representation of teeth crowning and tip relief modification. In order to study the effect of surface roughness on the gear pair power loss, the tooth root mean square (RMS) surface roughness for different crowning and tip relief modification are varied in the range 0.4\u20133.6 \u03bcm. The surface roughness of the current design of planetary hub gears of the JCB Max-Trac rear differential is 1.6 \u03bcm. Idealised spur gears with finite line contact geometry are very sensitive to misalignment and manufacturing errors which cause edge-loading of their contacts, leading to edge stress generated pressure spikes, similar to that of straight-edged rolling element bearings (Johns and Gohar, 1981; Kushwaha at al", + " Therefore, it can be concluded that the crowning amount falls in the range 50%\u201380% would be beneficial. At the beginning and the end of a meshing cycle with no tip relief, an impact and sharp rise in the generated contact pressure would ensue. In order to attenuate this effect, the involute profile in the tip region is relieved. An optimum length of tip relief region would enable smoother load variation from a pair of teeth to the next. The extent of tip relief in length and amount should be determined [see Figure 2(b)]. In order to study the effect of tip relief amount on the gear pair power loss, the amount of tip relief is changed from 25% to 150% of the current base design. Amounts of tip reliefs for different cases are listed in Table 2. A map of the results obtained is illustrated in Figure 7, showing the total power loss with different tip relief amount and surface roughness. Figure 8 represents the percentage change in the total power loss with respect to the current design. It shows that the overall power loss can be reduced by 12% in comparison with the current design", + " To further investigate the accuracy of the obtained optimum point, the optimisation steps are further reduced in its vicinity (i.e., 50% relief in steps of 5%). Figure 9 shows the percentage change in the total power loss with respect to the current base design for the various relieved amounts between 25%\u201375% of the existing design. Figure shows that the 50% tip relief amount is indeed the optimum case. In order to study the effect of change by the extent (length) of tip relief [specified as h in Figure 2(b)] on power loss, this is reduced from the current design (base value) by 25%. Table 3 shows the changes in the relief length for different cases. Figure 10 show the total power loss with different length of tip relief and surface roughness. Figure 11 shows the percentage change in the total power loss with respect to the current design. It shows that the power loss can be decreased by as much as 10%, representing a 50 W reduction per meshing cycle. Applying a 25% tip relief length yields the highest power loss reduction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003782_s12555-019-0436-3-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003782_s12555-019-0436-3-Figure2-1.png", + "caption": "Fig. 2. Landmark detection.", + "texts": [ + " The CV model can be described by the following MIMO model with a disturbance as follows:{ x\u0307 =Ax+Bu+\u03b5 , y =Cx, (16) A = [ L 0 I2 0 ] , B = [ \u039b 0 ] , C = [ 0 0 1 0 0 0 0 1 ] , I2 = [ 1 0 0 1 ] , L =\u2212(M\u0304T \u22121)\u22121V\u0304 T \u22121, \u039b = (M\u0304T \u22121)\u22121, l\u0307A =VA, \u03b8\u0307A = \u03c9A, (17) where A, B, and C are parameter matrices of the CV system, matrix T is defined by (3), u = [ \u03c4R \u03c4L ] \u2208 Rm (m = 2) is a torque input vector applied to the two wheels of the CV, x = [ VA \u03c9A lA \u03b8A ]T \u2208 Rn (n = 4) is a system state vector, lA is linear displacement and \u03b8A is an orientation of the CV, y = [ lA \u03b8A ]T \u2208 Rp (p = 2) is an output vector of the CV, and \u03b5 = [ \u03b51 \u03b52 \u03b53 \u03b54 ]T \u2208 Rn is a disturbance vector. The positioning of the CV is implemented by three steps: Landmark detection step using laser sensor Lidar, EKF prediction step based on encoders data, and EKF update step based on landmark positions. 3.1. Landmark detection step using laser sensor Lidar In this paper, laser sensor Lidar A-2 is used to detect landmarks. The detected landmarks using laser sensor Lidar can be expressed with polar coordinate as (di, \u03b2i) are shown in Fig. 2. In Fig. 2, OXY is a global coordinate frame, Cxy is a local coordinate frame of the CV, MxLyL is a local coordinate frame of laser sensor Lidar, \u03bb is a distance between C and M, \u03c1i is a distance between the CV and the ith landmark, \u03b1i is an angle between CV and the ith landmark current position, di is a distance between the sensor and the ith landmark, and \u03b2i is a scanning angle between sensor and the ith landmark. The ith landmark position in the local coordinate frame of the laser sensor Lidar can be written as follows: (xiL,yiL) = (di cos\u03b2i,di sin\u03b2i) for i = 1,2, \u00b7 \u00b7 \u00b7 ,n" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001568_med.2013.6608768-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001568_med.2013.6608768-Figure4-1.png", + "caption": "Figure 4. Athena Vortex Lattice model of aircraft and lift distribution", + "texts": [ + " The geometric definition of the aircraft, the modeling of the 8 segmented control surfaces on each wing half, and the flight conditions are modeled using the Athena Vortex Lattice (AVL) software package. AVL utilizes the vortex lattice method to determine the aerodynamic loads along the span of all aerodynamic surfaces, interprets the geometric configuration of the aircraft, and discretizes the wing into a finite element mesh. The AVL aircraft model and the corresponding static lift distribution can be seen in Fig. 4. AVL allows the user to obtain individual lift forces for each element in the mesh. The total lift load on each segmented portion of the wing is calculated by summing the elemental lift forces that are located within the geometric bounds defined by the segmented control surfaces. Permutations for each control surface deflection are modeled at varying angles of deflection in order to quantify the lift contribution of each control surface on the aircraft. From these permutations and corresponding lift forces, it is possible to find the rate of change of lift per degree of deflection that each flap contributes to each of the segments along the span" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000033_icus48101.2019.8996014-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000033_icus48101.2019.8996014-Figure7-1.png", + "caption": "Fig. 7: Expand the line segment to rectangle.", + "texts": [ + " In the simulation environment, we regard obstacles as a circle, and the robot as a point. Considering that robot is not a point in the actual scene, so we expand the line segment into a rectangle to leave a safe space for the robot. Judging whether the line segment collides with obstacles or not equals judging the relative position of the rectangle and the circle. Detailed steps are as follows. Step 1: expand the line segment to rectangle. After obtaining line segment Li,i+k, expand Li,i+k towards its two sides to obtain rectangle ABCD (Fig. 7). The length of ATi+k depends on robot\u2019s size and we set ATi+k = 0.2. In Fig. 7, the point O is the midpoint of Li,i+k, and the coordinate of point O is{ XO = 1 2 (Xi +Xi+k); YO = 1 2 (Yi + Yi+k), (7) where (Xi, Yi) and (Xi+k, Yi+k) are the coordinates of point Ti and point Ti+k, respectively. The angle between line segment Li,i+k and the x-axis is \u03b81, and the angle between line OP and the x-axis is \u03b82 (Point P is the center point of the obstacle). Step 2: translation and rotation transformation. Translate the center point O of rectangle ABCD to the origin point and get a new point O" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003899_012009-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003899_012009-Figure2-1.png", + "caption": "Figure 2. Commonly Used Robot Calibration Methods", + "texts": [], + "surrounding_texts": [ + "RIACT 2020 IOP Conf. Series: Materials Science and Engineering 1012 (2021) 012009\nIOP Publishing doi:10.1088/1757-899X/1012/1/012009\n2. Should standard calibration quality be related to the repeatability of the calibration instrumentation and the method of calibration? 3. What would be the optimal ways in which robot measurement can be selected? 4. How is observability of robot kinematic error parameters related to the selection of calibration configurations and method? Requirements for possibilities of practical implementation of uniform calibration system: 1. Robot calibration improvisation for faster, efficient & inexpensive calibration is needed. 2. The calibration load needs to be decided either by manufacturer or by end user. 3. Designing a robot to accommodate on-line calibration capability. 4. Use of advanced technology to make robot calibration more simple and user friendly. With regards to the implementation issues, the \u201cfast and cheap\u201d guideline automatically rules out expensive instrumentation such as Laser Tracker system, Camera Calibration, Coordinate Measuring Machine or Theodolite. From the calibration cost viewpoint the use of above instrumentation is an extremely costly and time-consuming process.\n6. A Robot Calibration System\nThere are many factors to consider during the calibration solutions. For the actual industrial robot, measurements are necessary to determine the best fit set of parameters. Tools positions with respect to the joint positions are required for measurement. Since the fit is as good as the original data collections. That is the significance of getting the valid set of data.\nIt is well known that the data used in calibration require sufficient joint exercise. This means the joints\nmust move significantly. These must include repeatability, setup-time, type of measurement, contact vs. noncontact, required software, and controller interface and cost. However, set-up time and cost is the major decision driver. Many calibration systems are having high cost and less significance in the robotic field, many production industries and researchers still not used these concepts.\nIndustry trends currently heading towards macro to micro analysis. Depending upon the application, i.e. for many large scale applications such as pick and place, welding operation etc this approach has worked well. But at the same time it's costlier for smaller applications.\n6.1 Brief discussion on Various Calibration Measuring Techniques Currently industrial robots are still practice programming using teach pendant. Human operators are required to guide the robot to the desired locations. The motions are recorded in the controller and that can be edited easily if required, within the robotic language and played back to perform the repetitive tasks. The robot calibration phase is the most significant step for any industrial robot.\nMost of the calibration measurement instruments are available such as dial indicator, external ball bar system, camera calibration, laser tracker systems, theodolite, or coordinate measuring machines are too slow or overly expensive or both.\nA flowchart as shown in Fig. 1, is simply a graphical representation of steps for selection of calibration methods. It shows steps in sequential order and is used in presenting the flow processes selection. Cameras calibration and camera vision systems have a standard automation component.", + "RIACT 2020 IOP Conf. Series: Materials Science and Engineering 1012 (2021) 012009\nIOP Publishing doi:10.1088/1757-899X/1012/1/012009\nIntegrated multi-camera systems of a robot cell are used for inspection, part identification, and real time inspection of assembly accuracy.\nFew cameras may fix in the robot workspace area and the others may fix to the moving end effector to assist the online component alignment. This system requires additional hardware and the additional software.\nIn calibration of robot, a theodolite is used for automatic partial pose measurement system. It required a\nrotational platform measurement from the camera system for finding out the spherical illuminate target. After calibration, the system can continuously track the target with sufficient speed, and accuracy. Since this approach requires dedicated system hardware and only a partial pose can be measured.\nA frame is mounted on tool coordinator at the end effector centre. After driving each axis to its reach limit, the Coordinate Measuring Machine (CMM) position indicator has set to zero. As CMM is used for pose measurement in industrial robot calibration.", + "RIACT 2020 IOP Conf. Series: Materials Science and Engineering 1012 (2021) 012009\nIOP Publishing doi:10.1088/1757-899X/1012/1/012009\nKey equipment for calibration is measurement systems. Measurement equipment and calibration systems must be appropriately matched for reliable identification of model parameters. The above table 1 presents measurement systems. Non-contact measurement systems with respect to their use in robot calibration.\nStatic measurement based on theodolites and triangulation principles. Another way is dynamic approach which is based on a laser tracking interferometer used to measure distance and the direction to determine position and orientation.\n7. Conclusions After critically analysing all facts related to existing robot calibration methods, their comparative analysis, making a ready reference flowchart for selection of suitable method following conclusions have been summarised as below:\n1. This study highlights the importance of low cost industrial robot calibration that can be used as a\ncheck on robot manufacturing quality and maintenance needs. The calibration procedure must be industry usable. This means that the procedures should be automated. The goal is to develop low cost equipment that can easily be used on the shop floor for incidental and periodic robot calibration.\n2. Based on the review of reference article, it is observed that the measuring system is having an\nimportant issue for finding a robot\u2019s process parameters.\n3. It is observed that some proposed standards for better understanding of testing and various\nspecification of a robot process performance with reference to the industry. Testing should be carried out on various point to point and continuous path in the work space.\n4. Finally the data produced from a calibration procedure is useful not only for improving robot\naccuracy, but also for analysing the quality in which a robot, or several robots of a single type, is manufactured. Because the calibration data is very extensive, a system is being developed to manage and analyse them based on the needs of the calibration procedure and the needs for quality control.\nReferences\n[1] B.W. Mooring, Z.S. Roth, R. D. Morris, Fundamentals of Manipulator Calibration, Wiley, New\nYork, (1991), 301-305\n[2] B. Greenway. Tutorial robot accuracy, Industrial Robot: An International Journal, 27(4) (2000),\n257- 265.\n[3] ABB ROBOTICS, Product specification \u2013 IRB1520 Robot documentation M2012, rev J, RW5.13,\n(2012), pp11-15, 12-17.\n[4] ISO 9283 Manipulating Industrial Robots-Performance criteria and Related Test Methods\nInternational Standards Organization (1998).\n[5] G. Boye, A. Olabia, J. Palosa, 3D metrology using a collaborative robot with a laser triangulation\nSensor. Procedia Manufacturing 11, (2017), 132-140.\n[6] M. Alia, M. Simic, F. Imad. Calibration method for articulated industrial robots. Procedia\nComputer Science 112, (2017), 1601\u20131610.\n[7] Z. Usman, P. Radmeh, An investigation of highly accurate and precise robotic hole measurements\nusing non-contact devices. International Journal of Metrology and Quality engineering 2016, 204- 214." + ] + }, + { + "image_filename": "designv11_34_0002608_vppc.2016.7791617-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002608_vppc.2016.7791617-Figure3-1.png", + "caption": "Fig. 3. Flux distribution of the proposed CPM-MG.", + "texts": [ + " The pole arc coefficient is the ratio of the width of each magnet to the pole pitch of the corresponding rotor. In the SPM-MG, the pole arc coefficient of both inner and outer PM are 0.5, therefore, the N-pole and S-pole magnets fully occupy the SPM space. However, in the CPM-MG, the pole arc coefficient of inner and outer PMs are 0.8 and 0.75 only, therefore, the PM volume is reduced, compared with that of the SPM-MG, to 75% on the outer rotor and to 80% on the inner rotor, respectively. B. Flux Distribution Fig. 3 shows the flux distribution of the proposed CPMMG when the outer rotor and inner rotor are aligned with each other. As it can be seen, the salient ferromagnetic iron pieces of the rotor together with the adjacent N magnets provide flux a path, that is, the iron pieces replace the S magnets in the conventional SPM-MG. Due to the much wider magnet, the flux tends to concentrate in the salient ferromagnetic iron pieces and the saturation in these areas becomes an issue when adopting analytical methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000656_s106836661405016x-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000656_s106836661405016x-Figure3-1.png", + "caption": "Fig. 3. (a) Loading, (b) experimental, and (c) calculated trajectories of shaft motion in connecting rod bearing in engine by Mit subishi Motor Company.", + "texts": [ + " This is explained by small changes of zone where condition p > 0 is ful filled, which leads to a decrease in the number of iter ations needed to solve the hydrodynamic contact problem, except for the first step at an initially unknown contact configuration. The algorithm was tested using the modeled FB, for which the trajectory of the motion of the shaft in the connecting rod bearing of an automotive engine by the Mitsubishi Motor Company was measured [7]. It can be seen in Figs. 3b, 3c that trajectories have sim ilar characteristic regions. At shaft rotation angles cor responding to steady loading (Fig. 3a), stabilization regarding eccentricity is observed. Variations in the relative eccentricity can be observed upon changing the loading. Along with other testing calculations, these results allow one to draw a conclusion about the JOURNAL OF FRICTION AND WEAR Vol. 35 No. 5 2014 HYDRODYNAMIC ANALYSIS OF FRICTION BEARINGS 399 applicability of the described algorithms for hydrody namic calculations of real slide bearings. SOFTWARE IMPLEMENTATION Developed methodology for calculating hydrody namic characteristics of bearings of gas turbine com pressor GTK10 I, crosshead bearings of marine die sel, high pressure pump station UNP55 250, opposite compressor and other aggregates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003802_s12541-020-00438-1-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003802_s12541-020-00438-1-Figure12-1.png", + "caption": "Fig. 12 Relationship among global, pelvis-fixed and foot-fixed coordinate frames", + "texts": [ + " The actual foot vertical force Fact z and the actual roll/pitch moments Mact Roll ,Mact Pitch were derived by converting the vertical force and moments in the foot-fixed coordinate frame obtained from the 6-axis force/torque into those in the pelvis-fixed coordinate frame. In Eq.\u00a0(13). R0 global is a rotation matrix of the global coordinate frame with respect to the pelvis-fixed coordinate frame, R0 foot is a rotation matrix of the foot-fixed coordinate frame with respect to the pelvis-fixed coordinate frame, and Fsensor is the force vector measured from the 6-axis force/torque sensor (see Fig.\u00a012). Equation\u00a0(14) shows the PI controllers to control the vertical force and roll/pitch moments at each foot. Kp,Fz , Kp,Mp , Kp,Mr are the proportional gains and Ki,Fz , Ki,Mp , KiMr are the integral gains of the vertical force controller and pitch/roll moment controllers, respectively. The total vertical force and roll/ pitch moments of each foot from the feed forward controller and feedback controller are finally derived as Eq.\u00a0(15). fupper = [ F QP z,L,u F QP z,R,u M QP Pitch,L,u M QP Pitch,R,u M QP Roll,L,u M QP Roll,R,u ]T (13) joint = JT a [ Ftotal x Ftotal y Ftotal z Mtotal Yaw Mtotal Pitch Mtotal Roll ]T On the other hand, reference horizontal position trajectory required for foot position control is derived by the Preview control method [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure4-1.png", + "caption": "Fig. 4 Critical contact situation between cutter and workpiece", + "texts": [ + " From which we can see, the cutting depth z and cutting arc length l of IEi vary with the rotation of cutter. For any IEi of cutter blade, as long as the cutting angle \u03b8 at every moment and its corresponding cutting depth z are confirmed, its swept area si can be calculated by Eq. (5). Based on swept areas of different IEi along cutting edge, cutter wear situation can be analyzed qualitatively. Figure 2 is just a sketch of cutting region; there are several different cases in actual machining. It is necessary to analyze each case separately. Figure 4 shows one case of contact situation between cutter and workpiece. Figure 4a is the whole vision, while Fig. 4b shows the enlarged cutting region. Point O1 is the center of the workpiece, point A is the intersection of two cutter locations p1 and p2, C is the outermost point of cutting edge, P is the tangency point between cutting edge and workpiece, offset e is the horizontal distance between axis of cutter and axis of workpiece, ap is the cutting allowance, and \u03b1 is the angle generated by cutter axis of two adjacent cutter locations (i.e., the cutter rotates angle \u03b1 relatively to the workpiece during one period from former blade cutting done until the current blade cutting done, see Eq. (25) for more details). The cutting region is the shadow area on YOZ plane whose vertexes are A, B, and C. Based on Fig. 4, keep all other factors constant, if offset e decreases to e2 in Fig. 5, the marginal parts C-C2 of cutting edge will not take part in machining. While offset e increases to e1, the peripheral cutting edge C1-C1\u2032 will take part in machining. So Fig. 4 shows the critical case of offset e. According to parameters R, Rc, and ap and geometric relationship in Fig. 4a, the critical offset ec is, ec \u00bc Rc\u2212 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2\u2212 R\u2212ap 2q \u00bc Rc\u2212 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2Rap\u2212ap2 q \u00f06\u00de Figure 6a shows the cutting region with maximal offset emax, when former cutter margin just reaches point A, as the red lines. O Z YP A ap O1 \u03b1 2 P' emax R-ap (a) Geometric model with emax O PA A' e=emax e>emax missed cutting region cutting region when e>emax (b) Missed cutting region when e>emax Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001400_detc2013-12646-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001400_detc2013-12646-Figure2-1.png", + "caption": "FIGURE 2. FORCE DIAGRAM FOR THE PLATFORM OF THE CABLEDRIVEN ROBOT", + "texts": [ + " In this section moment-resisting capability of the cable robot is investigated during the translational motion and the effect of the constant magnitude of pre-tension on this capability is discussed. The constant pre-tensions in passive pairs are necessary in order to have a constant orientation for Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/02/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2013 by ASME the moving platform. The end-effector will be in static equilibrium if force and moment equilibrium equations are satisfied. The force diagram of the robot is illustrated in Fig. 2, W is the gravity force and m1, m2 and m3 stand for 3 motors. Writing moment equilibrium equations about point C, static equilibrium equations may be obtained as: 0e CG W JT Wr r W (1) Where, 1 2 3 4 5 6 7 8 9 T T T T T T T T T T T (2) 6 9 |a pJ J J (3) 1 2 3 3 1 3 1 3 1 6 3 0 0 0a u u u J (4a) 3 31 2 2 1 6 54 5 4 6 6 54 5 4 6 6 6 p Cc CcCc Cc Cc Cc u uu u u u J r u r ur u r u r u r u (4b) In Eqn. (1), J is Jacobian matrix, re is the external wrench (expressed at point C) and CG is the position vector from C to center of gravity G. In Eqn. (2), Ti is the tension in ith cable as illustrated in Fig. 2 and in Eqn. (4b), is the position vector from C to ci. During positioning operation since the pre-tension in each pair has a constant magnitude, three constraints exist which express that the sum of tensions in two cables of each pair must remain constant. However, the tensions in two cables of each pair have not the same magnitude necessarily. Having 6 static equilibrium equations from Eqn. (1) and three constraints, we can obtain the tensions in all 9 cables. We may combine Eqn. (1) and three constraints and write: 6 1 3 1 3 1 0 0 0 e t CG t t W Wr P J T r W P P (5) where Pt is the constant magnitude of pre-tension and J\u2019 is: 9 9 0 0 0 1 0 0 1 0 0 0 0 0 0 1 0 0 1 0 0 0 0 0 0 1 0 0 1 pa JJ J (6) J\u2019 is a square 9\u00d79 matrix, so we can rewrite Eqn" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000318_j.finel.2015.07.008-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000318_j.finel.2015.07.008-Figure5-1.png", + "caption": "Fig. 5. The dynamic model of gear mesh effect.", + "texts": [ + " (7), where the tooth section is simplified to the combination of the rectangular section and the trapezoid section. kg \u00bc Fmesh \u03b4Br\u00fe\u03b4Bt\u00fe\u03b4s\u00fe\u03b4G \u00f07\u00de where Fmesh is the mesh force acting on the tooth; \u03b4Br is the corresponding bending deformation of rectangular section; \u03b4Bt is the corresponding bending deformation of trapezoid section; \u03b4s is the corresponding shear deformation; and \u03b4G is the corresponding elasticity deformation caused by the root tilting. Assume i and j are the nodes of the gear pair, where i is on the driving pinion, which is shown in Fig. 5. \u03c8 ij is defined as the angle between the tangent plane of the gear contact surfaces and the axis-z and described as \u03c8 ij \u00bc \u03d5ij\u00fe\u03b1ij driving gear rotate counterclockwise \u03d5ij\u00fe\u03b1ij \u03c0 driving gear rotate clockwise ( \u00f08\u00de where \u03b1ij is the pressure angle of the gear pair and \u03c6ij is the angle between the centerline of the gear pair and the axis-y (Fig. 5). The helix angle \u03b2ij is defined as positive when with a righthand helix driving gear, whereas negative when with a right-hand helix driving gear. The dynamic equations of the nodes i and j connected by the spring-damper element for gear mesh without considering the damping effect can be described as K ijuij \u00bc Fm \u00f09\u00de where Fm is the force vector of the meshing; uij is the displacement vector of nodes i and j, and defined as u\u00bc \u00bdui; vi;wi;\u03b8xi;\u03b8yi;\u03b8zi;uj; vj;wj;\u03b8xj;\u03b8yj;\u03b8zj T \u00f010\u00de The mesh stiffness matrix of the gear pair K ij in global coordinate system with 12th order can be described as K ij \u00bc kij UaT ij Uaij \u00bc k1;1 \u2026 k1;12 \u22ee \u22f1 \u22ee k12;1 \u22ef k12;12 0 B@ 1 CA \u00f011\u00de where aij \u00bc \u00bd sin \u03b2ij sin \u03c8 ij cos \u03b2ij cos \u03c8 ij cos \u03b2ij sgnri cos \u03b2ij sgnri sin \u03c8 ij sin \u03b2ij sgnri cos \u03c8 ij sin \u03b2ij sin \u03b2ij sin \u03c8 ij cos \u03b2ij cos \u03c8 ij cos \u03b2ij sgnrj cos \u03b2ij sgnrj sin \u03c8 ij sin \u03b2ij sgnrj cos \u03c8 ij sin \u03b2ij The symbol sgn is defined as sgn\u00bc 1 driving gear rotate counterclockwise 1 driving gear rotate clockwise ( \u00f012\u00de As early as 1980s, many scholars have carried out extensive researches in the dynamic model of TPJB, and many simplified models of TPJB are provided [17,18]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003564_icra40945.2020.9196986-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003564_icra40945.2020.9196986-Figure4-1.png", + "caption": "Fig. 4. The round peg (left) and aluminum profile with fixture (right) that are used for the experimental evaluation.", + "texts": [ + " Note that no adaption of the impedance is performed, and that the robot uses the same controller parameters throughout the entire insertion. A possible extension of the presented work is to adapt the controller parameters according to the current sample distribution. The setup consists of a KUKA LBR iiwa robot with 7-DOF and joint torque sensors for the measurement of external forces and impedance control. We investigate the approach in two scenarios: a round peg-in-hole and the insertion of an aluminum profile into a fixture (Fig. 4). Geometric properties and initial uncertainties are summarized in Table I. The algorithm is implemented in a partially paralleled manner, with 16 threads for collision checking, so that a command rate of up to \u22485 Hz can be realized. By default, we use a set of N = 320 samples. Further parameters of the observation algorithm are chosen according to [1]. In our current implementation, the KUKA RoboticsAPI is used to send the commands at the rate of our observation algorithm. The low level impedance controller of the robot runs at a higher controller rate >1 kHz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001928_1350650115579677-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001928_1350650115579677-Figure6-1.png", + "caption": "Figure 6. Assembly of the wheel and dynamic balance rings.", + "texts": [ + " The whole assembly is mounted on a 30mm thick steel table, which in turn is fixed on a heavy concrete bed, as shown in Figure 5, providing a solid support for the rig and eliminating the effect of other sources of vibration. The table is supported on the concrete bed via four adjustable cushions which enables the height of the table adjustable and the table horizontal. There are two balancing rings bolted to the wheels to balance the rotor. The balance weights are fixed in the balance rings, as illustrated in Figure 6. The rotor is balanced via a biplane dynamic balance method.9 A programme of rotor dynamic balance is designed based on a LabVIEW software, which detects the amplitude and the phase of the imbalance and directs the balancing of the rotor by decreasing the couple imbalance through moving the balance blocks. The rotor is driven by an electric motor, whose rotational speed can be infinitely varied from 0 to 3000 r/ min by a frequency control unit. While the rotational speed of the rotor may be infinitely varied from 0 to 6000 r/min via a pair of belt pulleys whose drive ratio is 2, as shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002946_j.precisioneng.2020.03.001-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002946_j.precisioneng.2020.03.001-Figure4-1.png", + "caption": "Fig. 4. Definition of the physical quantities and variables.", + "texts": [ + " The deformation of the highest cusp in this case is called the minimum contact deformation dc. Even if the contact surface has large waviness, the highest cusp is guaranteed to be deformed to the minimum contact deformation dc or more. Therefore, if the waviness of the contact surface is smaller than the minimum deformation dc (w < dc), real contact can be realized at all cutter-mark crossings regardless of the condition of the contact surface. Here, the minimum contact deformation dc is calculated for a surface finished using a tool with an R-shaped cutting edge. Fig. 4 illustrates the definitions of physical quantities and variables required for the calculation of dc. From geometric conditions, the following equations hold. h\u00bcR ffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 l2 4 r (4) x\u00bcRsin\u00f0\u03b1\u00fe \u03b8\u00f0x\u00de\u00de Rsin\u03b1 (5) sin\u03b1\u00bcR h R (6) y\u00f0x\u00de\u00bc l 2Rsin \ufffd\u03c0 2 \u00f0\u03b8\u00f0x\u00de \u00fe \u03b1\u00de \ufffd (7) where R is the nose radius of a cutting tool; l is the pick feed; and h is the cusp height. Fig. 5 shows a schematic of the real contact area after applying the CMC method. The upper and lower cusps deform by the same amount because both cusps have the same shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001186_s12541-013-0101-3-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001186_s12541-013-0101-3-Figure4-1.png", + "caption": "Fig. 4 Experimental set-up in floating test to estimate stability of porous bearing with bushing irradiated with Nd:YAG laser", + "texts": [ + " The volumetric flow rate of air q is measured for each disc by a flowmeter and the supply Table 1 irradiating conditions in surface modification of porous bushings Averaging power Q 205 (W) Pulse frequency f 400 (s-1) Duty ratio 80 (%) Diameter of spot d 1 (mm) Scanning velocity V 3 (m/min) Scanning pitch \u0394L 0.20, 0.30, 0.35 (mm) Flow rate of shield gas (Ar) 0.03 (m3/min) Fig. 1 Experimental set-up in laser surface modification for porous bushing Fig. 2 Change in apperance of porous bushing caused by laser irradiation pressure P1 and the delivery pressure P2 are also measured to obtain the differential pressure \u0394P = P1 - P2. Then, surface modified discs are used as the porous bushings of thrust bearing employed for the stability test as shown in Fig. 4. A shaft and a weight are floated by the air supplied via the porous bushing, and the floating height, the gap between the shaft and the bearing surface h, is obtained by means of averaging values measured with three displacement sensors. Figure 5 shows an experimental set-up to investigate load capacity and stiffness of the bearing which the surface modified disc is used as the bushing for. In this test, a load cell is mounted on the top of shaft to measure the axial force generated by the supply pressure of air", + "30 (mm), the heat input into the porous bushing is increased and the fluid resistance is increased to 7.27 (m3MPa/min). Thus, flow of the air through the porous bushing can be controlled by laser irradiating with various conditions. Figures 7 (a), (b) and (c) are fluctuation of floating height (= gap) for three kinds of bushings which are measured in the floating test with a constant load of 40 (N). In each figure, three curves represent output signals of displacement sensors which are mounted on the shaft floating as shown in the experimental set-up, Fig. 4. Figure 7(a) is obtained for the porous bushing as sintered. The height of shaft seems stable at low supply pressure of 0.04 and 0.07 (MPa), but it causes self-induced vibration14,15 when the pressure reaches 0.075 (MPa), slightly higher than a critical pressure. Meanwhile, the proposed surface modification with laser increases the critical pressure. In Fig. 7(b), vibration of floating shaft is caused for the supply pressure higher than 0.15 (MPa). This is because of the increase in fluid resistance which is caused by the laser surface modification as shown in Fig", + " And then, decreasing the pitch between scanning paths in laser surface modification process, the re-solidified layer is developed on the bushing surface and the fluid resistance is increased furthermore. In Fig. 7(c), therefore, the vibration is not caused even at high pressure 0.6 (MPa) for the thrust bearing with high resistance bushing. Figure 8 shows the change of gap between the shaft and the bearing surface h associated with the supply pressure P1 in the floating test. The floating gap is obtained by averaging values measured by three displacement sensors shown in Fig. 4. As can be seen, the gaps gradually increase as the supply pressure becomes high. For the bushing as sintered, and the bushing whose surface are modified with relatively large scanning pitch of laser, self-induced vibration is caused. However, the surface modification with small scanning pitch improves the dynamic stability of thrust bearing, so that the higher pressure air can be supplied without undesirable vibration of shaft. Figures 9(a) and (b) show the load capacity of thrust bearing W measured for the different floating gaps h in the stiffness test (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001276_s12541-014-0425-7-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001276_s12541-014-0425-7-Figure10-1.png", + "caption": "Fig. 10 The bead angle", + "texts": [ + " Each bead shape has its own ratio on the bead width and undercut depth. When R1 and R4 values are known, preference of the bead shape can be discovered. The numerical model to identify preference of the bead shape is as follows. Bead shape (6) Bead shape (7) Bead shape (8) In this process, 5 input parameters and 3 output variables are used to model the relationships between inputs and outputs. The input parameters include the laser power, the gap between two parts, the tensile load, and the bead angle (refer to Fig. 10). The structure of the BPNN consists of 4 nodes in the input layer, 12 and 7 nodes in the first and the second hidden layers, and 3 nodes in the output layer. The training data of the model include the laser power (kW), the gap between two parts (mm), the tensile load (MPa), the bead angle (\u00b0), the bead height (mm), the undercut depth (mm), and the bead shape. The performance of the BPNN for process number 1, 2, 4, and 7 are expressed in terms of the root mean square error (RMSE). The RMSE converged on 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003363_j.mechmachtheory.2020.104027-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003363_j.mechmachtheory.2020.104027-Figure7-1.png", + "caption": "Fig. 7. CVT chain entering pulley.", + "texts": [ + " 6 , the chain behavior that causes this vibration exciting force can be assumed to be the elastic deformation of the pin and fluctuation in the chain chordal velocity caused by the chordal vibration. This research focused on the fluctuation in the chain chordal velocity and examined a method to reduce it. Based on past research results, an effective means of reducing fluctuation in the chain chordal velocity is to shorten pitch l p [4] . However, if pitch l p is too small, the strength of the link, which is a component of the chain, will decrease. Therefore, this study examined a method to change the noise performance of the chain without adjusting pitch l p . Fig. 7 is a schematic diagram of an actual chain entering a pulley. The units in the chain chord are shown in yellow and the units in the pulley are shown in gray. The contact points C k between the pins are indicated by the blue circles and the contact points of the pins with the pulley are indicated by the red circles. The chain tension acts on the contact points C k between the pins. In the units in the chain chord, the contact points C k between the pins are aligned in a straight line due to the chain tension" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003825_j.synthmet.2020.116662-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003825_j.synthmet.2020.116662-Figure1-1.png", + "caption": "Fig. 1. (a) Optical image, (b) cross-sectional view (schematic), (c) 3D view (schematic) of spin-coated PANI-SSA film on glass over silver electrodes.", + "texts": [ + " Pre-defined 80 nm thick, 3 mm wide silver (Ag) electrodes were thermally evaporated on the glass substrate maintaining 2 mm spacing between two adjacent electrodes. The dispersion of PANI-SSA (2 wt%) in DMSO was sonicated and stirred prior to coating on top of Ag electrodes using spin coating method with spin speed 1250 rpm for 60 s. The samples were then annealed at 70 \u25e6C for 2 h to evaporate the solvent completely. Thickness of PANI-SSA films were measured using non-contact profilometer (Bruker, Contour GT). The optical image and schematics of the planar device are shown in Fig. 1. Humidity sensors based on HCl doped PANI (PANI-HCl) with similar configuration as that of PANI-SSA were also fabricated for comparison. The current-voltage (I-V) characteristics of the device were performed by measuring current upon applying a voltage sweep from \u2212 5 V to + 5 V using Keithley 2400 Source Measure Unit (SMU). Humidity sensing experiments were carried out in-situ using the Memmert CTC-256 humidity chamber having relative humidity (RH) range 10\u201398%, temperature range from \u2212 40 \u25e6C to 180 \u25e6C with temperature ramp of 10 K/min and a source measurement unit (SMU, model: Keithley 2400) for monitoring I-V characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003367_s11432-019-2671-y-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003367_s11432-019-2671-y-Figure3-1.png", + "caption": "Figure 3 (Color online) Diagram showing our LOS implementation.", + "texts": [ + " Generally, with sufficient propulsion and a suitable lift-to-drag ratio, such a robot can easily dive with a stable pitch angle. Thus, we can shift the control target from the depth to the pitch angle. Another benefit of controlling the pitch angle is that the tracking path becomes smoother, since the robot\u2019s body attitude cannot change sharply. In principle, depth control is a two-dimensional issue, and can be simplified for identifying tracking points at the desired depth. We can then employ LOS guidance to map the desired tracking points to pitch angles. As shown in Figure 3, given the target depth d, we consider a vector \u21c0 a that is perpendicular to zg. Then, taking the robot\u2019s centroid as the center, we draw a circle with radius R that intersects the vector \u21c0 a at the points A and B. Then, we select the point B as the target point. Based on the real-time depth z, we can then obtain the target pitch angle as \u03b8d = arctan ( de \u2016 \u21c0 a\u2016 ) , (5) where { de = d\u2212 z, d2e + \u2016 \u21c0 a\u20162 = R2. Here, \u2016 \u00b7 \u2016 indicates the Euclidean norm. 4.2 Control system 4.2.1 CPG model In this paper, we use a CPG model to control the robotic dolphin\u2019s swimming mode, as this can effectively ensure the control signals produce smooth motion transitions, even in the face of sudden parameter changes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000344_gt2015-43161-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000344_gt2015-43161-Figure2-1.png", + "caption": "Fig. 2 Experimental apparatus for measuring the structural dynamic characteristics of the proposed bearing", + "texts": [ + " Table 1 shows the principal dimensions of the rotor, the bearing bush, and the straight spring wires used in calculations. Dynamic characteristics of the proposed flexible structure To solve the threshold of instability of the proposed bearing numerically, the dynamic stiffness kw and damping coefficient bw of the proposed flexible structure made of straight spring wires are required. In this paper, therefore, the dynamic characteristics of the proposed flexible structure were experimentally obtained by using the frequency-response method. Figure 2 shows the experimental apparatus for measuring the dynamic characteristics of the proposed flexible structure. The flexible structure is inserted in the housing, and supports the fixed shaft instead of the bearing bush. The fixed shaft has the same diameter as the bearing bush. The piezoelectric actuator for oscillating the housing is located at the left side of the housing, and the piezoelectric load cell for measuring the dynamic force generated by the piezoelectric actuator is sandwiched and attached between the piezoelectric actuator and the housing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001040_s00170-013-5169-5-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001040_s00170-013-5169-5-Figure4-1.png", + "caption": "Fig. 4 Reference systems of the parallel kinematic mechanism", + "texts": [ + " X 0 \u00bc g Q\u00f0 \u00de \u00f02\u00de The D\u2013H method obtains the nonlinear equation system to model the mechanism. This method uses four parameters (distances di, ai, and angles \u03b8i, \u03b1i) to model the coordinate transformation between successive reference systems. The homogenous transformation matrix between frame i and i\u22121 depends on these four parameters (see Eq. (3). i\u22121A i \u00bc cos\u03b8i \u2212cos\u03b1i\u22c5sin\u03b8i sin\u03b1i\u22c5sin\u03b8i ai\u22c5cos\u03b8i sin\u03b8i cos\u03b1i\u22c5cos\u03b8i \u2212sin\u03b1i\u22c5cos\u03b8i ai\u22c5sin\u03b8i 0 sin\u03b1i cos\u03b1i di 0 0 0 1 2 664 3 775 \u00f03\u00de The reference systems used to calculate the transformation matrices are shown in Fig. 4. Equations (4) and (5) model SPS chains. B AT 1\u00bcB 0T \u22c5 0 3T \u22c5 3 4T \u22c5 4 7T \u22c5 7 AT \u00f04\u00de B AT 3\u00bcB 10T \u22c5 10 13T \u22c5 13 14T \u22c5 14 17T \u22c5 17 A T \u00f05\u00de Equation (6) models the chain 2. B AT 2\u00bcB 8T \u22c5 8 9T \u22c5 9 AT \u00f06\u00de In these equations, the superscript i, denotes the analyzed chain and B ATi represents the matrix that expresses the moving platform coordinates in the base reference system (SRB). Equation (7) shows the transformation matrix that enables us to obtain the moving platform coordinates in the reference system B by means of Euler angles ZYZ", + " The geometric parameter initial values must be estimated, and they should be close to the real values to ensure the algorithm convergence. Table 1 shows the D\u2013H initial parameters used to solve the kinematic model. The index i denotes the mechanism joint. (b) Calculation of the actuator elongation to place the moving platform in the desired pose from the Euler parameters, the anchorage positions and the current encoder readings. The elongation of the actuators can be calculated as the Euclidean distance between chains 1 and 3 (see Fig. 4) in the reference system B as shown in Eqs. (8) and (9). L1 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x7\u2212x0\u00f0 \u00de2 \u00fe y7\u2212y0\u00f0 \u00de2 \u00fe z7\u2212z0\u00f0 \u00de2 q \u00f08\u00de L2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x17\u2212x10\u00f0 \u00de2 \u00fe y17\u2212y10\u00f0 \u00de2 \u00fe z17\u2212z10\u00f0 \u00de2 q \u00f09\u00de The control algorithm performs the actuator movement in every analyzed position and the sensor readings, L1_encoder and L2_encoder, are stored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002006_s11071-016-2788-z-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002006_s11071-016-2788-z-Figure1-1.png", + "caption": "Fig. 1 Model-scaled helicopter", + "texts": [ + " The basic structure of the main rotor with Bell\u2013 Hiller stabilizing bar is introduced in Sect. 2. Flapping dynamics in hovering flight modes is derived and discussed theoretically in Sect. 3. The result in hovering flightmode is extended to vertical flightmode inSect. 4. A brief analysis in non-rotating frame is presented in Sect. 5. Simulation and experimental results in hovering mode are displayed in Sect. 6. This paper is concluded in Sect. 7. 2 The mechanical structure of the main rotor As is shown in Fig. 1, the Thunder Tiger Raptor 50 is a levo-rotational helicopter with two hingeless blade. Figure2 shows the azimuth angle\u03c8 which is measured from the downstream position of the blade. Being pivoted to the main rotor shaft, the Bell\u2013 Hiller stabilizing bar is a steel stick equipped with two paddles. Tilt of the swashplate, as is illustrated in Fig. 3, directly maneuvers the cyclic pitch of the paddles, which in advance contributes to a cyclic flapping motion of the Bell\u2013Hiller stabilizing bar. Since the stabilizing bar is a single rigid body, no collective flapping exists" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002017_1.4033364-Figure21-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002017_1.4033364-Figure21-1.png", + "caption": "Fig. 21 Roller\u2013raceway model", + "texts": [ + " It can be observed that for upper row rollers, moment and axial load act in the same direction, so higher roller loads are obtained in the upper row. 021503-8 / Vol. 139, MARCH 2017 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/935662/ on 01/31/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4.4 Friction Identification Analysis\u2014ADAMS Analysis. As mentioned earlier, in order to identify friction force during motion, a multibody-dynamic simulation program called ADAMS is used. In ADAMS, a basic raceway and roller model is generated (Fig. 21). Roller model is taken from three-dimensional drawings of the bearing. Raceways are generated as boxes while keeping raceway widths same as the bearing raceway. Both raceways and rollers are modeled as made of steel. Boundary conditions used in the model are given below: Lower box is fixed to the ground. Contact force is defined between roller and raceways. Force is applied from upper box perpendicular to upper box surface. For the upper and lower boxes to move parallel to each other, a parallel joint primitive is defined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002249_acc.2016.7525043-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002249_acc.2016.7525043-Figure1-1.png", + "caption": "Fig. 1. Platform support vessel: Frames definition and illustration of the lever arm from IMU to a GNSS PosRef", + "texts": [ + " To acquire high quality estimates of the attitude, one may employ the angular velocities from the IMU, together with measurements collected from e.g. an accelerometer or compass in a sensor fusion scheme, with or without aiding from position or velocity measurements. Examples of attitude observers applying nonlinear theory are [9]-[17]. Another sensor that is required for DP vessels is the vertical reference unit (VRU). Due to a DP vessel\u2019s design the respective PosRefs are located far from the nominal center of reference where the IMU may be located, as shown in Fig. 1. A distance between 30-50 meters is not uncommon. Therefore, the VRU is used to map the respective PosRefs\u2019 measurements from the measurement\u2019s location to a given point of interest utilizing roll and pitch readings provided by 978-1-4673-8682-1/$31.00 \u00a92016 AACC 985 a VRU. Since the distance between the PosRef and the point of reference can be quite large, it is most important that the roll and pitch signals from the VRU is accurate. The approach of this paper is clearly beneficial since no assumptions on the PosRef location on the craft are needed unlike the results of [18], [12], [14], [15] where the position reference is implicitly assumed located in the same point as the IMU", + " The lever arm between the IMU and given position reference (PosRef) system is taken into account, in contrast to previous results on nonlinear observers with USGES. Hence, the VRU is functionally embedded in the observer design, using the IMU. \u2022 Fault detection of PosRef position measurements with slowly emerging and varying faults using faultdiagnosis techniques and the observer structure posed. This paper employs two coordinate frames, North, East, Down (NED) and BODY, denoted {n} and {b}, respectively as seen in Fig. 1. NED is a local Earth-fixed frame, while the BODY frame is fixed to the vessel. The origin of {b} is defined at the nominal center of gravity of the vessel. The x-axis is directed from aft to fore, the y-axis is directed to starboard and the z-axis points downwards. A notation similar to the one in [19] is used: \u2022 pn - position of the vessel with respect to {n} expressed in {n} \u2022 vn - linear velocity of the vessel with respect to {n} expressed in {n} \u2022 \u03c8 - yaw angle between between {b} and {n} \u2022 \u03c9bb/n - body-fixed angular velocity \u2022 bbg - gyro bias \u2022 qnb - a unit quaternion representing the rotation between {b} to {n}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003656_s10015-020-00655-x-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003656_s10015-020-00655-x-Figure4-1.png", + "caption": "Fig. 4 Visualization of possible orientation of the rotor around Y P i and X P i", + "texts": [ + " For the expression of the model in a convenient way, the coordinate system of the i-th coaxial rotor ( i = 1, 2 ) is shown in Fig.\u00a03. The tiltable coaxial rotor-fixed coordinate systems lie in a plane and are separated by the angle of (see Fig.\u00a02), which are given by FPi \u2236 {OPi ;XPi , YPi , ZPi } (see Fig.\u00a03). Tilt angles (1)WRB = RZ( )RY ( )RX( ) i and i denote the ith coaxial rotor tilt angles about XPi and YPi , respectively. The initial tilt angles i and i are set to 0. The possible change about the tilt angles of the tiltable coaxial rotor is shown in Fig.\u00a04. In each tiltable coaxial rotor, the 360\u25e6 tilt mechanism is attached to the body coordinate system so as to change the pitch angle i of the coaxial rotor in 360\u25e6 (see Fig.\u00a04a), additionally, the 180\u25e6 tilt mechanism is connected to the coaxial rotors so that the roll angle i of the rotor can be changed in 180\u25e6 (see Fig.\u00a04b). The range of i is set to [ \u2212 \u22152, \u22152 ] and the range of i is set to [\u2212 , ] . Figure\u00a04a shows the tilt angle i around the YPi -axis of the rotor coordinate system, whereas Fig.\u00a04b shows the tilt angle i around the XPi -axis of the rotor coordinate system. The rotation order of the tiltable coaxial rotor that con- nected to the body coordinate system is from YPi - to XPi -axis in the coordinate system of the rotor. The rotation matrix BRPi from FPi to FB is written as follows: where, RZ , RY and RX are the rotations around the axis ZPi , YPi and XPi , respectively. The position vector OPi of the ith tiltable coaxial rotor in FB is defined as: The angular velocity Pi and its acceleration \u0307Pi of the i-th tiltable coaxial rotor can be obtained as follows: (2)BRPi = RZ((i \u2212 1) )RY ( i ) RX ( i ) (3)BOPi = RZ((i \u2212 1) )[l 0 h]T (4) { P1 = [" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003813_ffe.13393-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003813_ffe.13393-Figure7-1.png", + "caption": "FIGURE 7 Max stress of inclusion domain under different meshing status", + "texts": [ + " Therefore, the friction coefficient is specified as 0.05 in the present study. Considering that the variations of gear meshing position have a significant influence on the stress distribution near the inclusion in gears, the static contact analysis of gear pair with inclusion is presented for different meshing status. The cross section of cuboid inclusion is 0.299 \u00d7 0.123 mm2 with length of 2.35 mm. It is thought that after every clockwise rotation angle of 2 , the intermediate gear shaft would mesh with the low-speed gear. As shown in Figure 7, there are 13 rotation angles selected for the representation of different meshing status (i.e. mesh in or mesh out process of the gear teeth with oxide inclusion). It is clear that the max stress of inclusion domain varies with the different meshing status. For instance, the value of max stress achieves 445.99 MPa corresponding to the rotation angle of 12 . At this meshing status, there is a three-teeth meshing situation as shown in Figure 8, whereas the oxide inclusion is just below the contact position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003882_0954410020976611-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003882_0954410020976611-Figure2-1.png", + "caption": "Figure 2. Vector graph of aerodynamic forces. Figure 3. Vectored thrust frame and its deflection angle.", + "texts": [ + " Let U \u00bc FT sT , and then equation (4) can be rewritten as _x2\u00f0t\u00de \u00bc f2 \u00fe g2U\u00f0t\u00de (5) where f 2 \u00bc M 1 a FA \u00fe Fk \u00fe Fw \u00fe FGB sA \u00fe sk \u00fe sw \u00fe sGB ! \u00bc \u00f0f 2V\u00de3 1 \u00f0f 2x\u00de3 1 \" # ; g2 \u00bc M 1 a \u00bc g2V\u00f0 \u00de3 3 g2Vx\u00f0 \u00de3 3 g2xV\u00f0 \u00de3 3 g2x\u00f0 \u00de3 3 \" # Airship aerodynamic forces and moments on the right hand side of equation (4) can be expressed as follows12\u201315 FA \u00bc qSref Chcosu Chsinu Cz T sA \u00bc qSref lref Cmycosu Cmysinu 0 T 8< : (6) where u denotes the angle between the body x-axis and the horizontal component of the airspeed, see Figure 2. Sref and lref denote the reference area and length of the airship respectively, satisfying Sref \u00bc V 2=3 , lref \u00bc V 1=3 , V is the airship volume. q 2 R is dynamic pressure, Ch, CZ and Cmy 2 R are aerodynamic force coefficients along horizontal and z-axis, and moment coefficient along y- axis, respectively. a and b 2 R denote angle of attack and side-slip angle respectively. The wind-induced force vector of the airship is Fw \u00bc MBatw \u00fe x MBatw sw \u00bc JBa _xw\u00fex JBaxw ( (7) where tw and xw denote the linear speed and angular rate vectors of the wind" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001862_icrom.2014.6990963-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001862_icrom.2014.6990963-Figure5-1.png", + "caption": "Fig. 5. The model of bipedal robot.", + "texts": [ + " 9(a) and (b) represents the COM trajectories in different masses for fixed and moving ZMP, respectively. Comparing these figures, it can be concluded that the COM has less variation and stable motion in moving ZMP. For more comparison, the mean square errors (MSE) for different masses with fixed and moving ZMP are shown in Table. 1. It can be seen that MSE for moving ZMP is less than fixed one. To compare the joint trajectories for fixed and moving ZMP COM reference, the inverse kinematics problem is solved for 6-DOF bipedal robot shown in Fig. 5. The joint trajectories are obtained as shown in Fig. 10. Simulations are done for li = 0.3m (i = 1,2,\u20265) and lab = 0.1m. As illustrated in Fig. 10, the joint trajectories corresponding to moving ZMP have the same pattern with fixed ZMP but with less variation in joints. To compare the torque needed in joints for walking in both fixed and moving ZMP cases, the results obtained using Working Model 4D are shown in Fig. 11. It is seen from Fig .11 that the joint torques in moving ZMP is decreased. So it can be stated that, the energy consumption for biped locomotion with trajectories generated for moving ZMP is less than the one generated for fixed ZMP and it has optimal gaits" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure2-1.png", + "caption": "Figure 2. The relationship of shaper cutter and non-circular gear.", + "texts": [ + "2 r\u00f0 1\u00de \u00bc a\u00f01 e2\u00de 1 e cos\u00f0n1 1\u00de 8< : \u00f01\u00de According to the theorem of meshing tooth surface, the transmission ratio of curve-face gear pair can be shown as follows i12 \u00bc !1 !2 \u00bc R r\u00f0 1\u00de \u00f02\u00de And the rotating angle of curve-face gear 2 in the process of transmission can be shown as 2 \u00bc Z 1 0 1=i12d \u00bc 1 R Z 1 0 r\u00f0 1\u00ded \u00f03\u00de Based on the external generating method, the tooth surface of non-circular gear is founded by the involute cylindrical shaper cutter. The shaper cutter is rotating counter-clockwise and the non-circular gear is rotating clockwise in the process of meshing. Figure 2 shows the relationships of the positions and angles of the shaper cutter and non-circular gear. It is described in Figure 2 that the coordinate system rigidly linked with the frame of the shaper cutter is O1 X1Y1Z1 and the coordinate system rigidly linked with the shaper cutter is O01 X01Y 0 1Z 0 1. The coordinate system rigidly linked with the frame of the non-circular gear is O2 X2Y2Z2 and the coordinate system rigidly linked with the non-circular gear is O02 X02Y 0 2Z 0 2. Based on the coordinate transformation theory, during the meshing process, the coordinate transformation matrix M2010 from O01 X01Y 0 1Z 0 1 to O02 X02Y 0 2Z 0 2 can be established as follows M2010 \u00bc cos sin 0 L cos\u00f0 1 l\u00de sin cos 0 L sin\u00f0 1 l\u00de 0 0 1 0 0 0 0 1 2 664 3 775 \u00f04\u00de Matrix M220 is the coordinate transformation from O02 X02Y 0 2Z 0 2 to O2 X2Y2Z2 and can be established as follows M220 \u00bc cos\u00f0 1\u00de sin\u00f0 1\u00de 0 0 sin\u00f0 1\u00de cos\u00f0 1\u00de 0 0 0 0 1 0 0 0 0 1 2 664 3 775 \u00f05\u00de where r0\u00f0 \u00de \u00bc dr\u00f0 \u00de=d , L \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2\u00f0 1\u00de \u00fe r2s \u00fe 2rsr\u00f0 1\u00de sin p l \u00bc ar cos\u00f0L2 \u00fe r2\u00f0 1\u00de r 2 s=2Lr\u00f0 1\u00de\u00de, \u00bc 2 \u00fe 1 \u00bc arctan\u00f0r\u00f0 \u00de= r0\u00f0 \u00de\u00de. As shown in Figure 2, non-circular gear is rotating clockwise while shaper cutter is rotating counter-clockwise during the process of meshing, the at The University of Melbourne Libraries on June 5, 2016pic.sagepub.comDownloaded from generating surface 2 of non-circular gear can be expressed by the tooth surface of shaper cutter 1. Because the surface of non-circular is the conjugate tooth surface with shaper cutter, according to the space coordinate transformation theory, the mathematical equation of the tooth surface of non-circular gear 2 in coordinate O2 X2Y2Z2, can be established as follows r \u00f02\u00de 2 \u00f0 s, 1\u00de\u00bc rss\u00bdcos\u00f0 s os\u00de s sin\u00f0 s os\u00de \u00fe L cos\u00f0 1 l\u00de 8>>< >>: 9>>= >>; rss\u00bdsin\u00f0 s os\u00de \u00fe s cos\u00f0 s os\u00de \u00fe L sin\u00f0 1 l\u00de 8>>< >>: 9>>= >>; us 1 2 6666666666666664 3 7777777777777775 \u00f07\u00de The tooth surface of curve-face gear is founded by the involute cylindrical shaper cutter with the external generating method, the shaper cutter is rotating counter-clockwise and the curve-face gear is rotating clockwise in the process of meshing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.8-1.png", + "caption": "Figure 9.8 Third kind of position singularity of a Puma manipulator.", + "texts": [ + " (c) Third Kind of Position Singularity Equations (9.58) and (9.59) imply that the third kind of position singularity occurs if s\ud835\udf035 = 0, i.e. if \ud835\udf035 = 0 or \ud835\udf035 = \ud835\udf0b. In other words, this singularity occurs if the orientation of the end-effector is specified so that the approach vector is aligned with the front arm. However, while this singularity can occur easily with \ud835\udf035 = 0, it can hardly occur with \ud835\udf035 = \ud835\udf0b due to the physical shapes of the relevant links and joints. Therefore, this singularity is illustrated in Figure 9.8 with \ud835\udf035 = 0. If this singularity occurs, \ud835\udf034 and \ud835\udf036 become indefinite so that they cannot be found separately. This is because they turn out to be rotation angles about coincident axes, which are the axes of the fourth and sixth joints, and therefore they cannot be distinguished from each other. Nevertheless, their combination, i.e. \ud835\udf0346 = \ud835\udf034 + \ud835\udf036, can still be found as explained below. In the singularity with \ud835\udf035 = 0, Eq. (9.52) becomes eu\u03033\ud835\udf034 I\u0302eu\u03033\ud835\udf036 = eu\u03033\ud835\udf034 eu\u03033\ud835\udf036 = eu\u03033(\ud835\udf034+\ud835\udf036) = eu\u03033\ud835\udf0346 = C\u0302\u2217 (9", + " (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the wrist joint variables (\ud835\udf035, \ud835\udf034, \ud835\udf036). Therefore, it is the same as the third kind of multiplicity of the Puma manipulator. In other words, as illustrated in Figure 9.5, it occurs similarly as a wrist flip phenomenon without any visual distinction. A Stanford manipulator with d2 > 0 may have only one kind of position singularity. Equation (9.149) implies that it is the same as the third kind of position singularity of the Puma manipulator, which is explained in Section 9.1.5 and illustrated in Figure 9.8. (a) Angular Velocity of the End-Effector with Respect to the Base Frame Let \ud835\udf14 = \ud835\udf146. Then, Eq. (9.124) leads to the following expression. \ud835\udf14 = ?\u0307?1u3 + ?\u0307?2eu\u03033\ud835\udf031 u2 + ?\u0307?4eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 u3 + ?\u0307?5eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf034 u2 + ?\u0307?6eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 u3 (9.151) (b) Velocity of the Wrist Point with Respect to the Base Frame Let w = w(0) = vR = r\u0307. Then, w can be obtained from Eq. (9.127) as follows: w = (?\u0307?1eu\u03033\ud835\udf031 u\u03033)[u1(s3s\ud835\udf032) + u2d2 + u3(s3c\ud835\udf032)] + (eu\u03033\ud835\udf031)[u1(s\u03073s\ud835\udf032 + s3?\u0307?2c\ud835\udf032) + u3(s\u03073c\ud835\udf032 \u2212 s3", + " On the other hand, if the task to be executed necessitates to keep the manipulator in such a pose, then the freedom in \ud835\udf031 may be used to orient the folded arm links of the manipulator conveniently depending on the environmental conditions, e.g. for avoiding possible obstacles. (b) Second Kind of Position Singularity The second kind of position singularity of a Scara manipulator is the same as the third kind of position singularity of the Puma manipulator, which is explained in Section 9.1.3 and illustrated in Figure 9.8. Kinematic Analyses of Typical Serial Manipulators 279 (a) Angular Velocity of the End-Effector with Respect to the Base Frame Let \ud835\udf14 = \ud835\udf146. Then, Eq. (9.283) leads to the following expression. \ud835\udf14 = (?\u0307?12 \u2212 ?\u0307?4)u3 \u2212 ?\u0307?5eu\u03033(\ud835\udf0312\u2212\ud835\udf034)u2 \u2212 ?\u0307?6eu\u03033(\ud835\udf0312\u2212\ud835\udf034)e\u2212u\u03032\ud835\udf035 u3 (9.314) (b) Velocity of the Wrist Point with Respect to the Base Frame Let w = w(0) = vR = r\u0307. Then, w can be obtained from Eq. (9.288) as follows: w = \u2212u1(b1?\u0307?1s\ud835\udf031 + b2?\u0307?12s\ud835\udf0312) + u2(b1?\u0307?1c\ud835\udf031 + b2?\u0307?12c\ud835\udf0312) \u2212 u3s\u03073 (9.315) (c) Velocity of the Tip Point with Respect to the Base Frame Let v = v(0) = vP = p\u0307", + " Note that this singularity is nothing but the extreme version of the first kind of singularity considered in Part (a) of this section. Therefore, the arbitrariness of \ud835\udf033 can be resolved similarly. (c) Third Kind of Position Singularity Equations (9.703) and (9.704) imply that the third kind of position singularity occurs if s\ud835\udf037 = 0. Except the differences in the subscripts of the relevant angles, this singularity has the same appearance as the third kind of position singularity of the Puma manipulator, which is explained in Section 9.1.3 and illustrated in Figure 9.8. 328 Kinematics of General Spatial Mechanical Systems (a) Angular Velocity of the End-Effector with Respect to the Base Frame Let \ud835\udf14 = \ud835\udf146. Then, Eq. (9.641) leads to the following expression. \ud835\udf14 = ?\u0307?1u3 + ?\u0307?2eu\u03033\ud835\udf031 u2 + ?\u0307?3eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 u3 + ?\u0307?4eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 u2 + ?\u0307?5eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 eu\u03032\ud835\udf034 u3 + ?\u0307?6eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 eu\u03032\ud835\udf034 eu\u03033\ud835\udf035 u2 + ?\u0307?7eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 eu\u03032\ud835\udf034 eu\u03033\ud835\udf035 eu\u03032\ud835\udf036 u3 + ?\u0307?8eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 eu\u03032\ud835\udf034 eu\u03033\ud835\udf035 eu\u03032\ud835\udf036 eu\u03033\ud835\udf037 u2 (9.728) (b) Velocity of the Wrist Point with Respect to the Base Frame Let w = w(0) = vR = r\u0307" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000783_s1023193514110068-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000783_s1023193514110068-Figure2-1.png", + "caption": "Fig. 2. Cyclic polarization behavior of Ni/PDMA (SDS)/CPE in 0.1 M NaOH solution at anodic potentials at v = 50 mV s\u20131 (number of cycles = 5).", + "texts": [ + " After incorporating of Ni2+ ions into the polymer films, the polarization behavior was examined in 0.1 M NaOH solution using cyclic voltammetry tech nique. This technique allows the hydroxide film for mation in parallel to inspect the electrochemical reac tivity of the surface. Cyclic voltammograms of Ni/PDMA (SDS)/CPE at anodic potentials in the potential range of 0.1\u20130.75 V with a potential sweep 1022 RUSSIAN JOURNAL OF ELECTROCHEMISTRY Vol. 50 No. 11 2014 BANAFSHEH NOROUZI et al. rate of 50 mV s\u20131 are represented in Fig. 2. The current grows with the number of potential scans, indicating the progressive enrichment of the electroactive species Ni(II) and Ni(III) in the surface. The first positive potential scan generates a monotonically elevated cur rent flow giving a peak at a potential more positive than in the subsequent potential cycles. The peak shift is indicative of an overpotential required for the nucle ation and growth of [HOO(Ni/PDMA)]. The redox process of these modified electrodes is expressed as: Ni(OH)2 + OH\u2013 \u2194 NiOOH + H2O + e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002816_1350650119897471-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002816_1350650119897471-Figure8-1.png", + "caption": "Figure 8. Sketch map of the gear meshing process.", + "texts": [ + " cm is the meshing damping, cm \u00bc 2 ffiffiffiffiffiffiffiffiffi kmIe p , where Ie is the effective rotational inertia and is the damping ratio. e is the static transmission error of the gear pair. Lpi and Lgi are the friction moment arms of the ith meshing tooth pair for the pinion and gear and can be calculated as follows Lpi \u00bc AX\u00femod \u00f0rb1 1,Pb\u00de \u00fe \u00f0i 1\u00dePb \u00f011\u00de Lgi \u00bc AY mod \u00f0rb2 2,Pb\u00de \u00f0i 1\u00dePb \u00f012\u00de where Pb is the base pitch, the \u2018\u2018mod\u2019\u2019 function is calculated asmod x, y\u00f0 \u00de \u00bc x y floor\u00f0x=y\u00de, AX and AY are obtained as follows (on the basis of Figure 8) XY \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r21 r2b1 q \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r22 r2b2 q AY \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2a2 r2b2 q , AX \u00bc XY AY \u00f013\u00de where ra2 is the addendum radius of the gear. It should be pointed out that in actual gear meshing, there is a damping force produced by squeeze film effect, which will obviously affect the damping coefficient in the meshing process, and the change of meshing damping coefficient will also have a significant impact on the dynamic characteristics of gears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003165_j.matpr.2020.04.439-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003165_j.matpr.2020.04.439-Figure1-1.png", + "caption": "Fig. 1.Schematic of gear test rig.", + "texts": [ + " %mass fraction of polymer = (Wp/ Wc) \u00d7 100 (1) where Wc\u2013 mass of composite, Wp\u2013 mass of polymer,Wf \u2013 mass of metal foam. The tensile and flexural tests were conducted as per ASTM D638 and ASTM D790 to evaluate mechanical properties of polymer composite before fabricating spur gears. The tensile and flexural tests were conducted through universal testing machine \u2013 Tinius Olsen 25KT having maximum load capacity of 25 kN. Three samples were tested in each case and fourth sample was tested when not observing repeatability. Fig. 1 shows the schematic of the gear test developed in house. It has a 0.5 HP DC shunt motor with maximum speed up to 1500 rpm. It is coupled withan alternator having same capacity. The input to the motor and the field of alternator is controlled externally by two separate autotransformers. The autotransformer is used to adjustspeed and torque.There is an additional rectifier circuit consisting of diodes and capacitor for the AC to DC conversion. A 100 \u03a9 rheostat is used as the load across alternator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002591_s1068798x16120169-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002591_s1068798x16120169-Figure2-1.png", + "caption": "Fig. 2. Determining the gaps in engagement of the harmonic gear drive ahead of (a) and beyond (b) the pole.", + "texts": [ + " The azimuthal lateral gaps between the involute tooth profiles in the unloaded harmonic gear drive are determined after calculating the geometry of harmonic engagement and determining the auxiliary gear dimensions. The design of annular and disk harmonic generators ensures constant curvature of the median layer in the rim of the f lexible gear, within the engagement zones. The formulas for the lateral gaps are derived from the geometry of the involute harmonic engagement [9]. The gaps in the engagement may be calculated on the basis of Fig. 2. The distribution of the gaps within the engagement is determined by the position of the involute curves In1 and Ina describing the working profiles of the gears. The contact point of the involute curves may be within a section of the engagement ( ) \u221e = \u03d5\u03d5 = \u2212 \u2211 22 2 cos( ) ; 1p pS p ( ) \u221e = \u03d5\u03d5 = \u2212 \u2211 22 2 2 cos( ) ; 1p pV p p ( ) \u221e = \u03d5\u03d5 = \u2212 \u2211 22 2 sin( ) ; 1p pU p p ( ) \u221e = \u2212 \u03d5\u03d5 = \u2212 \u2211 22 2 sin( ) ; 1p pR p p ( ) \u221e = \u03d5\u03d5 = \u2212 \u2211 22 2 2 cos( ) . 1p pT p p RUSSIAN ENGINEERING RESEARCH Vol. 36 No. 12 2016 TORSIONAL RIGIDITY OF HARMONIC GEAR DRIVES 997 line (Fig. 2a) or outside it (Fig. 2b). In both cases, the azimuthal gap between the tooth profiles is , where is the central angle equal to the angular gap between involute curves In1 and Ina over the arbitrary circumference (radius rhy) of the hypothetical gear used in determining the azimuthal gap. As an illustration of the variation in the lateral gaps observed in the engagement of harmonic gear drives with Zf l = Z2 = 362, Zri = Z1 = 360, \u03b2 = 60\u00b0, x2 = 3, hc = 4.8 m, m = 0.5 mm we plot ji as a function of \u03d5 in Fig. 3. Vertical lines denote gaps at the onset of engagement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002095_s12206-015-1145-3-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002095_s12206-015-1145-3-Figure2-1.png", + "caption": "Fig. 2. Process of a peg-in-hole task and a hose assembly task.", + "texts": [], + "surrounding_texts": [ + "Peg-in-hole task strategies that make use of F/T sensors can largely be classified into four types." + ] + }, + { + "image_filename": "designv11_34_0003610_icuas48674.2020.9213956-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003610_icuas48674.2020.9213956-Figure5-1.png", + "caption": "Fig. 5: Continuous trajectory diagram.", + "texts": [ + " 5) Calculate the tangent points joining circles until the last one. 6) Calculate the tangent point between the last circle and the final point. This process is summarized in Algorithm 5. Based on previous data, a new arrange of points is given; this is, in addition to the initial point, final point, and waypoints, a new vector of tangent points is defined. With these complete data, continuous piecewise functions are defined; this through the use of lines and circles, parametrized by the path completion time, see Fig. 5. The piecewise functions are generated by the pseudocode in Algorithm 6. In order to validate the capabilities and performance of the overall algorithm, a series of tests is designed. Four test scenarios are proposed, where different arrange of obstacles are used to evaluate the trajectory generated; thus, it is concluded if the algorithm is capable of performing in a real 951 Authorized licensed use limited to: Middlesex University. Downloaded on October 18,2020 at 15:40:42 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure10-1.png", + "caption": "Fig. 10. (a) The Equivalent (von-Mises) stress in connecting rod; (b) Factor of safety of connecting rod.", + "texts": [ + " The objectives of Computer Aided Engineering (CAE) are:- To obtain accurate values of stresses and factor of safety so as to determine where the component is most likely to fail. To verify and validate the results obtained in the design calculations mentioned above. To obtain a simulation of the stresses as they occur in the component. The analysis of connecting rod was done by taking one end as a fixed support and a compressive force Fc \u00bc 4868:9N. The material taken was IS2062 Grade E250. The stresses developed is shown in Fig. 10 (a). From Fig. 10 (b), it can be seen that the connecting rod is safe to operate under the given conditions. Also, the results are similar to what was obtained through hand calculations in Section 2.2. The analysis of shaft was done by taking supports and forces as shown in Fig. 2 and Fig. 3. The material taken was AISI 4130. The results obtained are for gradual loading. The shear stresses developed is shown in Fig. 11 (a) and the factor of safety in Fig. 11 (b). It can be seen that the shaft will not fail under the given conditions and also, the results are similar to what was obtained through hand calculations in Section 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001859_2016-01-0814-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001859_2016-01-0814-Figure1-1.png", + "caption": "Figure 1. Strain washer installed under the spark plug. The cylinder head (10) have a threaded hole into which a spark plug 12 is screwed. A strain washer (15) is placed between the spark plug and the cylinder head. Part of a Figure from [18].", + "texts": [ + " Nevertheless, the data acquisition and analysis require expensive and highly specialized devices, due to the very high frequency content of acoustic emissions. The in-cylinder pressure curve can be obtained also by the measurement of the compressive forces acting on the cylinder head structure. The measurement devices can be mounted in various ways [15, 16, 17], but the most common location is beneath the spark-plug [18]. This setup requires the modification of the cylinder head, but the output of the strain washers is easily correlated to the in-cylinder pressure (Figure 1). The main advantage of this arrangement is that the loads on the spark plug caused by the pressure inside the cylinder are greater than those caused by extraneous loads (valve train dynamics, thermal loads, etc). For these reasons it is not required a complex post-processing of the data. A disadvantage of this setup is that the head of the engine has to be modified to house the strain washer. In addition, the spark ignition could cause electrical noises that can affect the accuracy of the incylinder pressure measurement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003389_acc45564.2020.9147655-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003389_acc45564.2020.9147655-Figure1-1.png", + "caption": "Fig. 1. CAD model of the dual system hybrid UAV airframe with the available control inputs marked in blue.", + "texts": [ + "00 \u00a92020 AACC 3858 Authorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on August 11,2020 at 22:52:08 UTC from IEEE Xplore. Restrictions apply. This section begins by introducing the modelling procedure of the hybrid UAV. Section II-B extends an fault-tolerant ISMC controller from the literature to model reference trajectory control. Sufficient conditions for the closed-loop stability are derived. In section II-C the HIL demonstrator and it\u2019s interfaces are presented in detail. Here the modelling of the dual system hybriv UAV is presented. Figure 1 shows the CAD model of the UAV, which has been introduced in [16]. The underlying configuration is a combination of a fixed-wing aircraft with the additional propulsion system of a multicopter. Such systems combine the ability to vertically take-off and land (VTOL) with the efficiency, the range and the velocity of a conventional fixed-wing aircraft. Here for VTOL maneuvers no tilting mechanism is needed like in similar tilt-wing or tilt-rotor configurations [22]. Furthermore, inherent overactuation properties are available, which are utilized in this work for FTC (see section III)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001296_01691864.2014.959051-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001296_01691864.2014.959051-Figure4-1.png", + "caption": "Figure 4. The workspace sphere to analyze the distributions of GR.", + "texts": [ + " Figure 3 shows the basic principle using the iterated approximating technique to search for them. To search an endpoint, we may select two points on the target pole, ensuring one outside but another inside the graspable region (points A and B in Figure 3, respectively), and then iteratively compute their mid-point and evaluate its possibility to being grasped. To analyze the distribution of the graspable region and to obtain two initial points for the approximating process, a simplified 3-D sphere representing the accessible workspace is used, as shown in Figure 4, with its boundaries:( W S \u2212 W C ) \u00b7 ( W S \u2212 W C ) = r2, (6) where W C = W B T [0 0 (l0 + l1)]T and r = \u2211n j=2 l j are the center and the radius of the spherical workspace, respectively, and li (i = 1, 2, . . . , n) is the length of the i th link and n is the degrees of freedom of the robot. D ow nl oa de d by [ M cM as te r U ni ve rs ity ] at 1 3: 06 0 9 Ja nu ar y 20 15 Table 1. The distribution of the graspable region. No. ABC Distribution of the graspable region 1 000 No solution or two segments lying on both sides of B 2 001 A segment on the right side of B, and the left side needs to be checked further 3 010 A segment across B 4 011 A segment across B including the whole right part (BC) 5 100 A segment on the left side of B, and the right side needs to be checked further 6 101 two segments distributed on both sides of B, both endpoints are valid 7 110 A segment across B including the whole left part (AB) 8 111 The whole segment AC is valid X B1 1M 2 1M 2 2M 3 1M 3 2M 3 3M 3 4M 2 XB i L In Figure 4, A and C are the two intersecting points of the sphere and the target pole, and B is their mid-point. Their coordinates may be obtained by solving the combined Equations (2) and (6). Note that utilizing an approximate spherical workspace, rather than the accurate workspace of the robot, is for convenience and simplicity in computation and analysis of the graspable region. Due to the complex expression of the accurate workspace, it is hard to obtain the initial points from the accurate workspace" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000814_ssd.2014.6808791-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000814_ssd.2014.6808791-Figure1-1.png", + "caption": "Fig. 1: Quadrotor with each rotor tilting about two axes", + "texts": [], + "surrounding_texts": [ + "Index Terms-Quadrotor, QRAV, helicopter, tilting rotor, De grees of freedom, VTOL\nI. INTRODUCTION\nQuadrotors have recently become a focus of research in Unmanned Aerial Vehicle (UAV) and flying robots applica tions. However, very little research, if any, targeted quadrotor as a manned aerial vehicle. Quadrotor Air Vehicles (QARV) may be employed in a wide range of commercial and military applications. Such applications may include: heavy transporta tion, construction of bridges and buildings, assembly of large pieces in factories, and rescue operations after natural disasters where roads and bridges are no longer usable. For military applications, QRAV may perform vertical takeoff and landing (VTOL) and can be used in manned operations for effective transport and for military deployment operations in hostile environments where VTOL is a requirement. Additionally, QRAV can have maneuverability that may be superior to helicopters, such as the APACHE helicopter. Conventional quadrotor is typically underactuated. It's composed of four fixed rotors which provide four input variables and has six degrees of freedom (DOF), 3 position and 3 orientations. The underactuated nature of typical quadrotors forces two translational motions to be coupled with two rotational ori entations, i.e. the x and y translation motions are coupled with the two rotational angles pitch and roll respectively. This coupling reduces the maneuverability and agility and limits tracking capabilities severely. For example, to move forward or sideways, roll or pitch angles is compromised and the UAV has to tilt. The UAV cannot go through tight openings and can't hover while having a tilted orientation. While these limitations are not of big impact on ordinary missions, critical missions demand much higher maneuverability. Conventional quadrotor modeling and control were covered in [1-4].\n978-1-4799-3866-7/14/$31.00 \u00a92014 IEEE\nMany breakthroughs emerged by researchers trying to over come the actuation difficulties in quadrotor. Tilt wing mecha nism was proposed in [5] and tilt rotor actuation in [6],[7]. In [5], a hybrid system was presented that has a tilt-wing mecha nism. The vehicle is capable of vertical takeoff/landing like a helicopter as well as flying horizontally like an airplane. Many researchers introduced techniques to decouple the motions of quadrotor. In order to maintain a zero net yaw moment, [8] proposed slightly slanted two opposite propellers with a small angle. It was shown that the four main movements : roll, pitch, yaw and heave can be completely separated using this design. In [9], a novel quadrotor design was presented. The four rotors were allowed to rotate about their axes W.r.t the main rotor body. This adds four extra inputs to have a total of eight inputs to the quadrotor. The design provides full actuation to the quadrotor position/orientation with two extra inputs.\nIn this paper, a novel quadrotor design is introduced. Each rotor is allowed to tilt around two axes W.r.t fixed body frame. The total number of inputs is increased to twelve. Although six inputs are enough to have a fully actuated system, twelve inputs may be required to impose arbitrary trajectories to more output independently. With this design, each of the twelve states (outputs) (6 positions/orientations - 6 transitional/rotational speeds) can be controlled independently and freely. More advantages and distinguished capabilities for critical missions are discussed later in this paper.\nThis paper is organized as follows: section II presents the dynamic model of quadrotor with two DOF's tilting propellers. Section III discusses the advantages of this design over conventional designs for manned operations and proves some of these advantages with simulation. Finally, the paper is concluded in section IV.\nII. SYSTEM MODEL\nThe model of the proposed design is introduced in our previous paper [10] but it's presented here for convention. In the following discussion we assume that the rotors are located at 01, O2, 03, and 04, and are tilted with respect to fixed rotor frames at these points Fig. I. The rotor frames are taken to be parallel to the body fixed reference frame at the center of gravity e.G. When the rotors are aligned along the body", + "z-axis, rotor 1 and rotor 2 are assumed to rotate counter-clock wise CCW, while rotor 3 and rotor 4 rotate clock-wise CWo The forward direction is taken arbitrary to be along the body x-axis.\nAssume that the rotational speed of the rotor i is given by Wi. Then we can say that the lifting thrust is given bw; and the drag moment is given by dw;. The orientation of the rotor is controlled by two rotations about the rotor fixed frame; ai, a rotation about the rotor y-axis, and f3i , about the rotor z-axis, as shown in Fig.2.\nTo find the forces and torques generated by each tilted rotor on the air vehicle , let R\ufffd: be the rotational matrix of the rotor with respect to fixed axis at Oi. Since the axes at Oi are parallel to the body axes at the center of gravity of the air vehicle, then [O\ufffd 0 G (Jisai] R7,' = R\ufffd = 0 S{JiSai o Gai (1)\n2\nThe thrust components of the ith rotor at the body CG. are then given by [0 0 G{JiSai] [ 0 ] Fi = 0 0 S{JiSai O2\no 0 Gai bWi (2)\nSimilarly, the moments of a titled rotor consist of two parts, the drag moment, and the moments generated by the thrust\ncomponents. These two components can be expressed as [ 0 0 Cf3isai] [ 0 ] Mi = 0 0 Sf3iSai 0 +ri x Pi\no 0 Cai dw;O(i) (3)\nWhere Cf3iSai == Cos(f3i)Sin(ai); 0 = [1,1, -1, -1] and ri is the vector from center of gravity to the reference point of the rotors, i.e.\nr1 = [l, 0, -h] r1 = [0, l, -h]\nr3 = [-1, 0, -h] rl = [0, -I, -h]\nh and I represent the vertical and horizontal displacements from the rotors to CG respectively.\nThe summation of forces acting on the quadrotor body is\nthen given by the dynamic equation:\n(4)\nWhere Xb is the position on x axis with respect to the body frame i.e. w.r.t the pilot. And REB is the body transformation matrix with respect to the earth inertial frame, and is given by\nREB = Rq, \u2022 Re \u2022 Rip [\ufffd: t: \ufffd] [C0 8 \ufffd S 0 8 ] [\ufffd \ufffd\ufffd _\ufffd\u00a2]\no 0 1 -S8 0 C8 0 S\ufffd C\u00a2 [CWC8 -SWC\ufffd + CwS8S\ufffd SWs\ufffd + CwS8C\ufffd ] SwC8 CWC\ufffd + swS8S\ufffd -CWS\ufffd + SWS8C\ufffd -S8 C8S\ufffd C8C\ufffd\n(5)\nW, 8, \ufffd are the body yaw, pitch, and roll respectively. Now, let Sl = [1>, B, \ufffd]T. The rotation dynamic equation is\nthen given by: I n = - (Sl x ISl ) - Me + M (6)\nWhere I is the inertia matrix of the quadrotor, Me is the gyroscopic forces, and is give by\n4 Me = L IR(Sl x wi)O(i) (7) i=1\nIRis the rotor moment of inertia. And", + "[\ufffd 0 Cfi,Sa,] [\ufffdJ Wi= 0 SfJiSai 0 Cai\n(8)\n4\nM= L Mi (9) i=l\nThe equations of motion can be easily placed in the form\nWhere\nx = f(X,U)\nx = [x, x, y, fj, Z, i, \u00a2,;p, e, e, cJ;, 1/;]\nNote that the combination of positions, velocities and ac\ncelerations are chosen to be convenient to the pilot.\nIII. CONTROL\nIn this proposed design, a control panel may be provided in order for a pilot or operator of a QRAV to access and manipulate a plurality of control parameters for each of the four rotors. The control panel includes inputs in the form of two joysticks and display screens. One of the joysticks may be used by the pilot or operator to control the forward speed x by moving the joystick forward and backward, and lateral speed y may be controlled by moving it horizontally left and right, while the forward acceleration X, or thrust, may be controlled by twisting the joystick. However, in highly maneuvering cases, as in the combat scenario, the pilot can switch the forward speed control to the forward acceleration control. In the acceleration control the forward acceleration x is proportional to the position of the joy stick. The neutral position of the joy stick could cause the aircraft to either hover or maintain its last forward speed. The second joystick may be used to control the rotational movements of the air vehicle. The forward/backward position may be used to control the pitch of the air vehicle e, the left/right positions may be used to control the roll \u00a2 of the air vehicle, while twisting the left joystick 2 may control the yaw angular velocity cJ;. The touch screen enables the pilot to limit the range of vehicle speed that can be reached by the full span of the joystick. For example in a pick-and-place mission to precisely install bridge construction parts, the range of speed control by the joystick can be limited to 1 or 2 meters per sec for precise motion and control of the air vehicle. Similarly, the pilot can set limits on the vehicle forward acceleration for specific missions. The set up can be saved and retrieved when the pilot starts similar missions. Possible mapping between pilot commands and quadrotor inputs is shown on the table in fig.3.\n3\nThe display screens may show information including one or more of: elevation, forward velocity, orientation of the air vehicle (roll, pitch, yaw), odometer, trip meter, fuel level, battery status, global positioning system (GPS) information, and geographic information system (GIS) information. The touchscreen displays and/or the display screens may show information including rotational speed of one or more of the four rotors, power consumption, and alarm status (temperature, overpower, overspeed, etc.).\nI V. ADVANTAGES AND SIMUL ATION TESTS\nThis proposed design can offer many advantages over all the existing designs in the literature. Some of these advantages are tested in this paper while other are left for future work. For example, not only the motions of quadrotor are decoupled, but also all the translational and rotational speeds can be controlled independent of the positions/orientations. This means that the quadrotor can move on a certain trajectory while maintaining specified speeds and orientations which gives this design su perior maneuverability. The free inputs can further be used to achieve additional tasks such as overcoming gust disturbances or even as brakes. On the other hand, while the additional inputs may be of great use during critical missions, they can be turned completely off when not needed to save power and reduce control complexity. In fact, the quadrotor is still fully actuated and the motions are completely decoupled using only any two opposite rotors. Failure of any of the rotors would not compromise the safety of the flight or behavior. Furthermore, if the rotors are allowed to rotate freely in a hemisphere, i.e.\na is allowed to reach proper angles; and motors are strong" + ] + }, + { + "image_filename": "designv11_34_0002440_10402004.2016.1245456-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002440_10402004.2016.1245456-Figure3-1.png", + "caption": "Fig. 3 \u2013 Thermal and structural boundary conditions scheme", + "texts": [ + " [9-12] need to be solved in both solids with its ACCEPTED MANUSCRIPT 10 respective boundary conditions taking into account the thermal deformations \u033f as shown in Eq. [12]. \u033f [9] \u033f \u033f ( )( ) \u033f ( \u033f ) [10] \u033f ( \u20d7 \u20d7 ) \u033f [11] \u033f [12] As boundary conditions for Eq. [9-12] it has been considered the external loads due to the interaction fluid-solids, and the interaction between both solids (asperity contact component), and the thermal expansion due to the temperature field. Only elastic deformations are considered and displacement is not allowed on the base of the support and the shaft edges as shown in Fig. 3 due to the high rigidity of the clamping elements. The asperity contact components aims to calculate the proportion of load carried by the asperities and by the lubricant under mixed lubrication. The approach has been applied and validated for line and point contacts by Liu (7) and Faraon and Schipper (8), but analytical equations were used for the fluid dynamics and solid deformations. The asperity contact component presented in this paper follows the same approach but it is implemented locally using CFD and FEA", + " If mixed lubrications occurs, a new source of heat per unit of area must be considered due to the friction between asperities (Bejan et al. (42)). The friction heat is distributed in the contact interface between both solids following the Charron\u2019s relation (43), shown in Eq. [30], where \u03b3j, is the proportion of heat generated in the interface transmitted to the journal. Kj,s, \u03c1j,s and Cp-j,s are the thermal conductivity, density and specific heat capacity of the journal and housing respectively. \u221a \u221a \u221a [30] Finally, the boundary conditions of the thermal phenomena are defined (see Fig. 3) as shown in Eq. [31], where refers to the symmetry plane domain, refers to the surfaces in contact with air at temperature , convection and solid conductivity , refers to isolated surfaces, refers to the film surface and represents the heat produced due to friction in the film surface. { \u20d7 \u20d7 ( ) \u20d7 \u0302 \u20d7 ( \u0302 ) \u222b [31] ACCEPTED MANUSCRIPT 16 The mixed-TEHL model is computationally expensive and the mesh should be optimized. Due to symmetry, only half of the bearing is meshed. The fluid dynamics phenomena are more critical than the solid mechanics of the bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003949_j.jterra.2020.11.004-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003949_j.jterra.2020.11.004-Figure5-1.png", + "caption": "Fig. 5. Each wheel for the experiment.", + "texts": [ + " Modulus Value Unit Name of parameters Reference c 762a \u00f0N=m2\u00de Soil cohesion Mizukami, 2013 C0 same as c \u00f0N=m2\u00de Soil-tool adhesion for a virtual plate Mizukami, 2013 cw 0 \u00f0N=m2\u00de Soil-wheel adhesion \u2013 g 9.81 (m=s2) Earth gravity \u2013 q \u2013 \u00f0N=m2\u00de Surcharge on the soil surface Decided by experiment d 15 \u00f0 \u00de External friction angle \u2013 b 90 \u00f0 \u00de Rake angle \u2013 c 1430 (kg=m3) Soil density Measured by experiment j 0.002\u20130.02 (m) Shear deformation modulus Mizukami, 2013 / 22:3a 32:5b \u00f0 \u00de Internal friction angle a: Mizukami, 2013, b: measured a: the value was decided by the shear test, b: the value was decided by the funnel test. and 80 mm in width (Fig. 5). The wheel mass is set from 1.0 to 2.5 kg every 0.5 kg. Tables 2 and 3 summarizes the experimental conditions and the soil parameters. The soil bin area is width, length, and high of 300; 1200, and 180\u2013200 mm, respectively, and Silica sand No. 5 fills in that area. In this system, the wheel can move to a vertical direction freely; that is, the towing motion does not affect a vertical motion. The wheel sinkage is set at 10\u2013 40 mm every 10 mm as the initial sinkage. The wheel unit connects to the towing unit with the rope that has low elasticity, and the rope tows the wheel unit through the towing motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000469_isie.2015.7281526-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000469_isie.2015.7281526-Figure1-1.png", + "caption": "Fig. 1. CDFIG: two DFIG connected back-to-back.", + "texts": [ + " Based on the experimental result of current control loop, the transient response of the CDFIG to current step changes in the Control Machine and the cross coupling disturbance are examined. In the power control loope, the decoupling of the active and reactive power is examined. The main contribution of this work was the construction of a simple and flexible test bench to validate various control strategies applied to a CDFIG. II. CDFIM MODEL The model of the CDFIG can be derived from the models of two DFIM connected back-to-back as show in figure 1. Disregarding magnetic saturation, the electrical machine model can be supposed linear and is defined by the following equations, considering the electrical dqO expressed in a general reference frame and the rotor connection in positive phase sequence[ 12] [13]: -g -g - R -,g d1jJr . ( 9 )-;:;;g Ur rZr + dt + J W - wep \ufffdr -g -g - R -,g d1jJsc . ( 9 )-;:;;g Usc scZsc + dt + J W - wepc \ufffdsc (1) (2) (3) The rotor connecting in the positive phase sequence makes the individual torque components operate in the same direction, but with the negative phase connection the individual torque work in opposite direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000826_s1052618814050136-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000826_s1052618814050136-Figure3-1.png", + "caption": "Fig. 3.", + "texts": [ + " 380 JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 43 No. 5 2014 NOSOVA et al. We define the number of degrees of freedom of the manipulating mechanism by the formula of A.P. Malyshev used for spatial mechanisms , (1) where, n is the number of elements; p5, the number of pairs of fifth class (one degree of freedom kine matic pair); and p4, the number of fourth class (two degrees of freedom kinematic pair). We use formula (1) for definition of the number of degrees of freedom of the mechanism with four degrees of freedom (Fig. 3). The spatial mechanism (Fig. 3) includes base 1, output element 2, working organ 3 interlocked with the output element 2 and three kinematic chains. Each chain contains a sliding motor\u20144, 4 ', 4 '', located parallel to one of the orthogonal reference axes; an initial rotary kinematic pair 5, 5 ', 5 ''; the initial element of the parallel crank mechanism 6, 6 ', 6 ''; rotary kinematic pairs of the par allel crank mechanism 7, 7 ', 7 ''; terminal rotary kinematic pair 8, 8 ', 8 ''; terminal element of the parallel crank mechanism 9, 9 ', 9 ''; as well as an intermediate element of the parallelogram and terminal element of the kinematic pair 10, 10 ', 10 ''", + " Let us calculate the number of degrees of freedom at this condition. In this case, we take into account the same elements that were used before, excluding interme diate elements of the parallelogram. Three sliding kinematic pairs will change twelve rotary kinematic pairs of three parallelograms. Substituting values in (1), we shall get W = 6 \u00d7 (13 \u2013 1) \u2013 5 \u00d7 14 = 2. This result is not true as well. Let us change the parallel crank mechanisms in two kinematic chains not containing a rotary motor (Fig. 3, I, II) by a universal joint (Fig. 3, Ia, IIa) and we shall consider the parallel crank mechanism as W 6 n 1\u2013( ) 5p5\u2013 4p4\u2013= W 3 n 1\u2013( ) 2p5 p4\u2013\u2013= JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 43 No. 5 2014 SYNTHESIS OF MECHANISMS OF PARALLEL STRUCTURE 381 the sliding kinematic pair in the third kinematic chain containing the rotary motor. Thus, altogether we get 15 elements and 16 one degree of freedom kinematic pairs in the mechanism. Substituting values in (1) we shall get W = 6 \u00d7 (15 \u2013 1) \u2013 5 \u00d7 16 = 4. This result is true: the number of degrees of freedom is four" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000413_med.2015.7158800-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000413_med.2015.7158800-Figure2-1.png", + "caption": "Fig. 2. 3DOF Hover dynamic system.", + "texts": [ + " The group formed by front and back motors (supply voltages 978-1-4799-9936-1/15/$31.00 \u00a92015 IEEE 520 given by vf and vb) causes the movement on the pitch and yaw axes, while the lateral motors (analogously vr and vl) move the roll and yaw axes [9]. The system has three encoders which measure the angular displacements in the three freedom axes of the plant from an initial position. Assuming a linear model (with the equilibrium point in which the propellers are aligned with the axes X , Y and Z, Fig. 2), the pitch movement can be described as: Jp \u22022p \u2202t2 = lKf (vf \u2212 vb) , (1) where, Jp is the equivalent moment of inertia about the pitch axis, p is the pitch angle, l is the distance from pivot to each motor and Kf is the propeller force-thrust constant [2]. In a similar manner, for the roll movement: Jr \u22022r \u2202t2 = lKf (vr \u2212 vl) , (2) in which, Jr is equivalent moment of inertia about the roll axis and r is the roll angle. The torques generated by the front and back propellers are called \u03c4r and \u03c4b, and similarly, the torques generated by the right and left propellers are \u03c4r and \u03c4l. As shown in Fig. 2, the torque generated by lateral propellers has a reverse direction compared to the torque generated by front and back propellants. The yaw movement is given by: Jy \u22022y \u2202t2 = \u03c4f + \u03c4b + \u03c4r + \u03c4l (3) Jy \u22022y \u2202t2 = Kt,c(vf + vb) + Kt,n (vr + vl) (4) in which Jy is the equivalent moment of inertia about the yaw axis, y is the yaw angle, Kt,n and Kt,c are torque constants that relate generated torque and voltage applied to the motor [2]. In this study the model from the manufacturer\u2019s manual is used [9] which uses the state vector, x(t)T = [y(t),p(t),r(t),y\u0307(t),p\u0307(t),r\u0307(t)] (5) and the control vector u(t)T = [vf (t), vb(t), vr(t), vl(t)]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure6.9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure6.9-1.png", + "caption": "Fig. 6.9 Numerical model where magnetic forces act on the rotor", + "texts": [ + " Literature about noise in electrical machines deals only with modes and frequencies of forces acting on the stator. Indeed the rotor-surface is coupled with the stator over the air-gap. The air of the air-gap is not capable to vanish in axial directions at frequencies of 500 Hz up to 10 kHz. Indeed the air is reacting like air in a piston air cylinder, [6.10, 6.11]. The phenomenon has two effects. On the one hand the rotor surface forces are coupled to the stator by the stiffness of the air-gap. On the other hand the machine has to be considered as a coupled structure between rotor and stator, see fig. 6.9. 98 6 Noise Based on Electromagnetic Sources in Case of Converter Operation The stiffness can be determined acc. equation 6.11: \u03b4 \u03c1 2 ' c C = , (6.11) with C\u2019: Stiffness per surface, c: sonic speed, \u03b4: air-gap; \u03c1 density of air. The assumption that the air-gap is a mere stiffness element is valid as long as the following relationship 6.12 is valid: 1<\u22c5\u03b4\u03c9 c , (6.12) with vibration angular frequency \u03c9. Larger frequencies or air-gaps would cause reflections. The stiffness of the airgap has been included in the numerical model acc. fig.6.9. The described measurement is a very special case. Forces, which act on the rotor, can only cause noise components, if the rotor and stator Eigen-frequencies of some modes are in the same range. If the Eigen-frequencies are different, the displacement amplitudes at the rotor surfaces are very small for an excitation with the stator Eigen-frequency. The forces, which act on the stator bore depend directly on the rotor surface displacement and will be quite small as well. 6.3 Eigen-Modes and Mechanical Calculation Methods 99 Force modes and mechanical Eigen-modes have to be differentiated strongly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003437_s00170-020-05886-7-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003437_s00170-020-05886-7-Figure18-1.png", + "caption": "Fig. 18 Images of the flexure mechanism measured by X-ray CT: a fabrication error analysis, b defect analysis, and c crosssectional view at the half-plane [10]", + "texts": [ + " There still remain significant gaps in fully understanding and establishing the relationship between the process, structure, property, and performance of AM parts. In this past research, the characteristics of crack and related challenges inherent to metallic parts fabricated via laser-assisted AM were analyzed by an X-ray computerized tomography (CT) scanner. The current CT scan software supplies extensive analysis tools for quantifying defects and deviations of wall thickness and creating nominal to actual comparisons. Figure 18a presents an example of the features that can be measured by using this technique. The right image shown in Fig. 18c is a cross section of the scan of a solid flexure and a solid model obtained using the CT scan data and the sectional planes are shown in Fig. 18b. Furthermore, the quality of the AM stage was investigated by the use of X-ray CT (METROTOM\u00ae 800, resolution ~ 5 \u03bcm). The scanned data was compared with the original CAD data, and the error of fabrication was analyzed. The variation (error of fabrication) over the whole flexure area was less than 0.1 mm (Fig. 18a), and the air-voids were not found in the flexure area as well as the ground frame (Fig. 18b). In addition, the geometry of the cross-sectional area at the half-plane along the Z-direction was monitored as seen in Fig. 18c and the obtained data shows that the designed stage was successfully fabricated without any air-voids and shape irregularity within the CT scanning resolution limit. We concluded that the X-ray CT scanning to be an alternative method for investigating the defect of AM-applied devices and the fabrication performance of the AM process. To obtain high precision motion quality, precision devices should have the ability to perform the repeatable motion and therefore it can be considered as a reliable motion mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000230_ijcaet.2017.086921-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000230_ijcaet.2017.086921-Figure2-1.png", + "caption": "Figure 2 Direction of excitation of spring", + "texts": [], + "surrounding_texts": [ + "An accelerometer is a device that measures proper acceleration which is not necessarily the coordinate acceleration, i.e., rate of change of velocity. Instead, the accelerometer sees the acceleration associated with the phenomenon of weight experienced by any test mass at rest in the frame of reference of the accelerometer device. Single and multi-axis models of accelerometer are available to detect magnitude and direction of the proper acceleration or g-force, as vector quantity, and can be used to sense orientation (because direction of weight changes), coordinate acceleration, vibration, shock, and falling in a resistive medium, a case where the proper acceleration changes, since it starts at zero, then increases. To acquire reading of acceleration for suspension system, accelerometer device is placed on the top surface of primary suspension and check the vibration or excitation of the springs in vertical, longitudinal and lateral directions with respect to time in millisecond. The excitations of springs are recorded during running of rail road vehicle for distance of about 60 km on actual Indian track condition at a maximum speed of 80 km/hr. From the readings, it is observed that the lateral vibrations of spring are comparatively larger than the longitudinal vibration is shown in Figure 3. It means the suspension spring bears a lateral deflection at the curvatures which is limited to 16 mm for middle axle springs because only middle axle are responsible to negotiate the curvature for rail road vehicle. The vertical excitations are extremely high and peak point is maximum between 20 sec to 30 sec as shown in Figure 4, indicates the suspension spring which bears the vertical load have vibrations in running condition which can cause failure of spring." + ] + }, + { + "image_filename": "designv11_34_0001107_icems.2014.7014032-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001107_icems.2014.7014032-Figure2-1.png", + "caption": "Fig. 2 Motor structure of original model (1/4 model). Fig. 3 FEA model to obtain the gap magnetic flux density by rotor coil (1/4 model).", + "texts": [ + " Then, the magnetic flux density distribution caused by magnets is reproduced by the rotor windings of the prototype to refer the air gap magnetic flux density. After that, no-load iron loss of the stator is confirmed whether it matches with the original one by FEA. No-load iron loss is generated by rotating magnetic flux which created by magnets. Therefore, it is considered that the iron loss can be reproduced by the rotating magnetic flux that made from a distributed current in the wound rotor even in the rotor is locked. Fig. 2 shows the quarter models of the investigated original motor. Nd-Fe-B permanent magnet is used in this machine. The stator has 24 slots with distributed windings. To examine the rotating magnetic flux distribution which created by wound rotor, the magnetic flux density in air gap created by a line current is basically analyzed. The air gap flux density is calculated by changing parameters of the coil current amplitude i, the distance of the coils from the center of rotor r, coil pitch \u03b8 and number of phases of coil" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002354_s40032-016-0339-5-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002354_s40032-016-0339-5-Figure3-1.png", + "caption": "Fig. 3 a Feature based slicing [5]. b Single slice in feature based slicing [5]", + "texts": [ + " A separate support structure has to be deposited to build overhanging structures [4]. Although, use of a distinct support material is quite common in nonmetallic AM processes, such as Fused Deposition Modelling (FDM), the same for metals is not yet available. Powder-bed techniques like Selective Laser Sintering (SLS) employ easily-breakable-scaffolds made of the same material to realize the overhangs. However, the same approach is not extendible to deposition processes like, LENS. A different approach to solve this problem is feature based slicing method as shown in Fig. 3. Unlike uniform and adaptive slicing techniques, where the thickness of a given slice is constant, in feature based slicing or inclined slicing. The thickness varies even within a given slice, based on its feature. Various approaches have been developed by researchers to identify the features in a CAD model and slice the part accordingly (either in parallel or inclined slices). Some of the investigators have used a recursive approach for creating the adaptive slicing of a part [6]. The CAD model is first sliced uniformly into slabs of thickness equal to the maximum available fabrication thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003312_metroind4.0iot48571.2020.9138245-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003312_metroind4.0iot48571.2020.9138245-Figure2-1.png", + "caption": "Figure 2. 3D model of the external frame with aluminum profiles and plywood.", + "texts": [ + " The development of the external frame has been realized taking into account the following constraints: (i) height and diameter of the robotic platform equal to 35 cm and 75 cm, respectively; (ii) height sufficient to host subjects during tests; and, (iii) easy to assemble and disassemble for maintenance operations. External frame has also been equipped with stairs to facilitate the access of subjects. The structure has been realized by using aluminum profiles for a total mass equal to 105 kg and dimension 2.6 m x 1.6 m x 1.4 m. Plywood was used to complete the structure. 3D CAD model of the external frame is shown in Figure 2. The previous version of the Roto.BiT3D is currently on service at two hospitals for research and clinical testing purposes: v1.0 at MARlab (OBG), and v2.0 at the 527 Authorized licensed use limited to: Carleton University. Downloaded on July 12,2020 at 18:20:19 UTC from IEEE Xplore. Restrictions apply. biomechanical laboratory of \u201cPolo Tecnologico\u201d (Don Gnocchi Foundation, Milan). The sensor system embeded in the Roto.BiT3D consisted in a load cells and a pressure matrix. B1. Load cells In order to evaluate the force and moment exerted by platform to the subject and vice versa, the robotic device is equipped with a uniaxial load cell on each arm and a 6-axis force/torque measurement device between the moving base and the main spherical joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002351_chicc.2016.7554080-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002351_chicc.2016.7554080-Figure1-1.png", + "caption": "Fig. 1: Frame definitions of 3D trajectory tracking", + "texts": [ + " Assuming that the gravity of the body and buoyancy are equal and their centers are located vertically on the z-axis of {B}, and ignoring the nonlinear hydrodynamic damping terms, roll motion and environmental disturbances, the kinetics of an under-actuated AUV equipped with a back propeller and one pair of stern control surfaces can be simplified into the following equation \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 u\u0307 = m22 m11 vr \u2212 m33 m11 wq \u2212 d11 m11 u+ \u03c4u m11 v\u0307 = \u2212m11 m22 ur \u2212 d22 m22 v w\u0307 = m11 m33 uq \u2212 d33 m33 w q\u0307 = m33 \u2212m11 m55 uw \u2212 d55 m55 q \u2212 mghs(\u03b8) m55 + \u03c4q m55 r\u0307 = m11 \u2212m22 m66 uv \u2212 d66 m66 r + \u03c4r m66 (2) where m(\u00b7) denote the system inertia coefficients, d(\u00b7) express hydrodynamic damping efforts, and h = zg \u2212 zb with zg and zb denoting the coordinates of the center of gravity and buoyancy. In addition, \u03c4u, \u03c4q and \u03c4r are the control inputs acting on the AUV in surge, pitch and sway directions, respectively. Hence, the AUV is under-actuated without control inputs in the sway, heave and roll directions, which leads to more degrees of freedom to be controlled than the number of the available control inputs. As depicted in Fig.1, an under-actuated AUV represented by Q tracks a desired 3D Cartesian trajectory with a specified timing law. Denote the inertial position and speed vector of the time-dependent particle P on the tra- jectory by pP (t) = [xP (t), yP (t), zP (t)] , and p\u0307P (t) = [x\u0307P (t), y\u0307P (t), z\u0307P (t)] , respectively. The size of its resultant speed vector is denoted by UP (t) = (p\u0307 P p\u0307P ) 1 2 . The problem of trajectory tracking control for an underactuated AUV can be formulated as follows: Given a desired 3D trajectory, the feedback control laws for the external force and torques acting on an under-actuated AUV are developed such that the distance from P to Q reduces to zero and the resultant speed of the AUV is identical with that of the particle P on the trajectory. In this section, the improved guidance law for 3D trajectory tracking control of an under-actuated AUV is presented, which is instrumental in designing a simplified feedback linearization PD controller to achieve position tracking as well as orientation tracking along the desired 3D trajectory. Associated with P , the corresponding trajectory frame {F} using the speed orientation of P as its x-axis can be built, as shown in Fig.1. To make the x-axis of frame {I} align with the x-axis of frame {F}, we only need two consecutive elementary rotations about the z-axis and y-axis of frame {I}. Hence, the y-axis of frame {F} is always parallel to the xy-plane of frame {I}, and its z-axis is defined by the right-hand rule. Define those two rotation angles \u03c7P and \u03c5P as follows \u23a7\u23a8 \u23a9 \u03c7P = atan2 (y\u0307P , x\u0307P ) \u03c5P = arctan ( \u2212z\u0307P\u221a x\u03072 P+y\u03072 P ) (3) The function atan2 is the arctangent function with two arguments and its range is (-\u03c0,\u03c0], while the range of the inverse trigonometric function arctan is (-\u03c02 ,\u03c02 )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure2-1.png", + "caption": "Figure 2. Mechanism of unit.", + "texts": [ + " An overall view of the traveling-wave-type omnidirectional wall-climbing robot is shown in Fig. 1, and its specifications are listed in Table I. The robot is composed of eight linked units, each of which connected by servomotors to enable rotation. This robot is essentially octagonal. TABLE I. ROBOT SPECIFICATIONS Maximum Diameter [mm] 650 Height [mm] 80 Weight [kg] 7.9 Control devise PMB-001 (Pirkus, Inc.) Power source Stabilized power supply 978-1-4673-9675-2/15/$31.00 \u00a9 2015 IEEE 2223 The structure and specifications of the units are shown in Fig. 2 and Table II, respectively. Mainly, each unit consists of two servomotors, a magnetic adhesion mechanism and a universal joint. One motor rotates the unit, and the other controls extension and magnet rotation. The servomotor for unit rotation is directly connected to the universal joint of the adjacent unit and controls the angle between the unit and its neighbor about the z-axis. Conversely, the servomotor for unit extension and magnet rotation is attached to the cam driver. The cam mechanism of the driver enables extension and contraction of the unit and magnet rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003088_tia.2020.2992579-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003088_tia.2020.2992579-Figure1-1.png", + "caption": "Fig. 1. Topology of PCLSPMV machine.", + "texts": [ + " Additionally, the starting and steady performances of PCLSPMV machine and DSLSPMV machine are validated by finite element analyzes (FEA) in Mengxuan Lin, Dawei Li*, Member, IEEE, Kangfu Xie, Xun Han, Xiang Ren and Ronghai Qu, Fellow, IEEE Comparison between Pole-Changing and DualStator Line-Start Permanent Magnet Vernier Machine P V Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 03,2020 at 06:44:11 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. section IV. Finally, experimental results will be given to validate the above analysis in section VI. II. STATOR WINDING POLE-CHANGING METHOD Fig. 1 shows the typical structure of LSPMV machine. The stator slot number Z, the rotor PM pole-pair number Pe and the armature winding pole-pair number Pa should satisfy: - a eP Z P (1) The pole-changing winding is adopted for stator winding. During starting process, the winding pole-pair number is changed by switches. In order to place the rotor starting unit on the periphery of the rotor, the interior PM rotor is adopted, which is different from the SPMV rotor invested in majority of papers. The wound winding is chosen as the rotor starting unit rather than squirrel cage for PCLSPMV machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure5.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure5.2-1.png", + "caption": "Fig. 5.2 The center of mass, and the relationship with any point p", + "texts": [ + "1 The Center of Mass 235 z-axis is chosen to be vertical. In particular, for navigation systems (aircrafts and boats) the positive direction of the vertical z-axis is defined pointing toward the center of the earth. The other two axes are chosen arbitrarily provided they fulfill the right-hand rule condition. The summation of the gravity effects of each particle should then be equal to the effect of the total mass at the center of gravity: \u222b B dp \u00d7 gdm = dc \u00d7 gm (5.7) where the inertial distances dp and dc (see Fig. 5.2) are the inertial position of the particle p and the inertial position of the center of mass cm respectively. Since the center of mass is a particular case, its location can be obtained from the rigid motion expression (3.39), i.e. dp = d + Rrp, where the relative position is the one of the center of mass: dc = d + R(\u03b8)rc (5.8) Using Eqs. (3.39) and (5.8), Eq. (5.7) can be written as (\u222b B ( d + Rrp ) dm \u2212 (d + Rrc) m ) \u00d7 g = 0 (5.9) Which after cancelling the identical terms becomes 236 5 Dynamics of a Rigid Body R (\u222b B rpdm \u2212 rcm ) \u00d7 g = 0 Whose non-trivial solution yields the definition of the center of mass as the weight average of the overall mass distribution over the body B: rc 1 m \u222b B rp dm (5.10) It is very common in the literature to express the center of mass of a body with respect to itself. This can be achieved in different ways. One of them is to redo the same analysis as before, but using the position of particles dp from the inertial position of the center of mass as (see Fig. 5.2): dp = dc + Rr (5.11) Then the non-trivial solution of (5.9) becomes: \u222b B rdm = 0 (5.12) This last equation establishes that the center of mass of a body is the point from which the distances r to every single particle on that body fulfils last relationship. Expression (5.12) can also be found from expression (5.10) if the center of the body\u2019s frame 1 is placed at the center of mass (see special case of Fig. 5.5), which is a practical use but not the general case. Then, Eqs. (5.10) and (5.12) are equivalents and both describe the place of the center of mass", + " 244 5 Dynamics of a Rigid Body Theorem 5.1 (3-D Parallel Axes Theorem) The inertia tensor of a body is augmented when the body rotates about an axe that does not intersect its center of mass. The augmentation is given by the semidefinite positive second moment of inertia(\u2212m[rc\u00d7]2 ) \u2265 0, where the vector rc = (rcx , rcy , rcz ) T \u2208 R 3 stands for the distance from the rotating point to the center of mass of the body, i.e.: I1 \u222b B [ rp\u00d7 ]T [ rp\u00d7 ] dm = \u2212 \u222b B [ rp\u00d7 ]2 dm = Ic \u2212 m[rc\u00d7]2 > 0 (5.26) Proof Notice after Fig. 5.2 (left) that rp = rc + r with a constant offset distance rc. Then the square product becomes [ rp\u00d7 ]2 = [(rc + r)\u00d7]2 = [rc\u00d7]2 + [rc\u00d7][r\u00d7] + [r\u00d7][rc\u00d7] + [r\u00d7]2. Notice that the crossed products vanish inside the integral after expression (5.12). Then expression (5.26) arise after the use of definition (5.23). Remarks \u2022 Note that the shifted Inertia Tensor I1 of a body is also a positive definite constant matrix due to property (1.99l) (i.e.: \u2212[a\u00d7]2 \u2265 0), with the following explicit expression I1 = \u23a1 \u23a2\u23a2\u23a3 Ixxc + m ( r2cy + r2cz ) Ixyc \u2212 mrcx rcy Ixzc \u2212 mrcx rcz Ixyc \u2212 mrcx rcy Iyyc + m ( r2cx + r2cz ) Iyzc \u2212 mrcy rcz Ixzc \u2212 mrcx rcz Iyzc \u2212 mrcy rcz Izzc + m ( r2cx + r2cy ) \u23a4 \u23a5\u23a5\u23a6 (5", + "57b) are not the same. However they are equivalent to each other if (5.57a) holds. Dynamics at the center of massNotice that the force f (0) happens to be the addition uniquely of the external forces over the body because the inner ones are mutually canceled due to Newton\u2019s 3rd law. Then, the force at the center of mass is a free vector: f (0) c = f (0) (5.58) f (1) c = f (1) (5.59) On the other hand, the torque as computed in (5.57b) is a line vector, with n1 defined by (5.52). Since r(0)p = r(0)c + r(0) (See Fig. 5.2, left), the torque n(0) 1 around the origin of frame 1 can also be expressed as n(0) 1 = \u222b B ( r(0)c + r(0) )\u00d7 d f (0) p = r(0)c \u00d7 f (0) + n(0) c (5.60) or equivalently with non-inertial coordinates: n(1) 1 = rc \u00d7 f (1) + n(1) c (5.61) 256 5 Dynamics of a Rigid Body where n(1) c is the torque around the center of mass cm, produced by al exogenous differential force acting on all and every particle of the body, expressed with noninertial coordinates: n(1) c = \u222b B ( r \u00d7 d f (1) ) (5.62) Then the inertial torque can be written as n(0) 0 = (d + r(0)c )\u00d7 f (0) + n(0) c Which comparing with expression (5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002017_1.4033364-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002017_1.4033364-Figure5-1.png", + "caption": "Fig. 5 Combined loading tests", + "texts": [], + "surrounding_texts": [ + "3.1 Test Procedure. Motion transfer diagram of the test assembly is given in Fig. 4. Corresponding locations of the friction identification tests are presented in the figure. In the scope of this work, five different groups of tests\u2014295 tests in total\u2014are conducted. First, in order to determine the internal friction of the test setup, constant speed torque values are measured for 13 different rotational speeds between 0.001 rad/s and 0.067 rad/s. In the second group of tests, the slewing bearing with its original configuration\u2014with upper and lower seals intact\u2014is tested for 13 different speeds and six different loading conditions. Third group of tests consists of measurement of the friction torque without upper seal in order to observe the effect of the upper seal. In the fourth group, the lower seal is taken off, too. By doing this, friction torque of the slewing bearing free from seals is obtained. In addition, the effect of lower seal on the friction torque is derived by using third and fourth group of tests. In the last group of tests, friction torque measurement of the slewing bearing under combined loading is carried out. These tests are performed without any seals in order to examine the effect of combined loading on the friction torque of the bearing. For each rotational speed and load combination, the bearing is driven with a constant speed in clockwise and counterclockwise directions. In Figs. 5 and 6, the test setup with tilting moment and axial loading is presented, respectively. The tilting moment is exerted to the bearing by hanging weights to the tip of a moment arm, while the axial load is applied by simply putting weights onto the bearing outer ring in vertical plane. 3.2 Internal Friction of Test Setup. Friction measurement test setup is used to measure friction torque of the slewing bearing. However, the test setup has an internal friction caused by the bearing of the test setup. Since the friction of this bearing is out of consideration for this work, this friction component should be eliminated. In order to measure internal friction of the test setup, it is mounted in the same orientation as slewing bearing friction measurement tests to simulate the same effect of orientation of the test setup. Since only the frictional torque of the test setup is of interest, pinion gear is not coupled with slewing bearing gear. In this way, slewing bearing load is eliminated during internal friction measurement of the test setup. However, influence of loading due to the gear mesh is ignored in these tests. The results of internal friction torque measurements are presented in Fig. 7. Internal friction of the test setup changes between 0.16 N m and 0.21 N m. The general trend of the curve is such that, when the rotational speed increases, the friction torque decreases, although there exist some discrepancies. Average Journal of Tribology MARCH 2017, Vol. 139 / 021503-3 Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/935662/ on 01/31/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use internal friction torque is calculated as 0.18 N m, which is, at some speeds, 14% of the measured friction torque values. So, it is vital to subtract internal friction torque from the total bearing torque. Since the rotation is transferred from the setup to the slewing bearing with a spur gear pair, a radial and moment loading is applied to the bearing unit of the setup. The radial force due to gear pair is calculated as Fr \u00bc Ft tan\u00f0a\u00de (1) where Ft \u00bc 2T d (2) In calculations, torque (T) is taken as the maximum torque measured in the tests which is 3 N m, diameter of the slewing bearing (d) is taken as 480 mm, and pressure angle (a) is taken as 20 deg. Then, the radial load exerted on the bearing unit is found to be 4.5 N. By multiplying the radial load with shaft length (0.1 m), the moment load exerted on the bearing unit is found to be 0.45 N m. Since the calculated loads are very small compared to load-carrying capacity of the bearing, the effects of radial and moment loadings are neglected in internal friction measurements of the test setup. 3.3 Friction Torque of the Slewing Bearing. As mentioned earlier, friction measurements are performed in different groups of tests with the aim of observing different frictional characteristics of bearing elements. When the test results are examined, the friction torque caused by the lower seal, by the upper seal, and by the bearing internal mechanism can be separated. In Fig. 8, effects of these three factors on friction torque are presented with respect to rotational speed in axial loading. There are two important issues which must be given attention. First, 65.5% of the total friction is caused by sealing operation while the rest which amounts to 34.5% is caused by bearing internal friction. That is, sealing of the bearing increased the friction torque about twice as much. Second, there is a major difference between the upper and lower seal friction torques. While the upper seal generates 10.7% of total bearing friction, the lower seal is responsible for 54.8% of total bearing friction, which is five times higher than that of the upper seal friction. This difference is caused by the sealing types. The upper seal is normal lip seal, since the upper part of the bearing is not exposed to harsh environmental effects. However, the lower seal is a lip seal with a garter spring in order to withstand harsh environments. In Fig. 9, the effects of the upper seal, lower seal, and bearing internal mechanism on friction torque with respect to changes in axial load are shown. At low loads, the lower seal friction is about 60% of the total bearing friction while at high loads this value decreases to 48%. Inversely, the bearing inner friction increases with load from 26% of total friction to 38%. The bearing inner friction is not affected by the load distribution at this level and it represents 14% of the total friction torque. 3.4 Bearing Inner Friction Torque in Axial Loading. Bearing inner friction torque arises from the resistance due to elastic hysteresis, deformation of the rollers and the raceways, sliding between bearing elements, and viscous friction torque. In Fig. 10, the bearing inner friction torque with respect to rotational speed is given. The bearing inner friction torque can be investigated in two regions: rotational speeds below and above 0.005 rad/s. For the rotational speeds higher than 0.005 rad/s, the friction torque increases with increases in speed. This trend changes for the rotational speeds below 0.005 rad/s. It looks that there is a dependence on load in this region. Up to 2000 N of axial load, the friction torque increases and then decreases slightly causing a local maximum at low speeds. However, for higher loads, the general trend of friction curves is decreasing, producing a local minimum. Dependence of bearing inner friction on load is presented in Fig. 11. Up to 800 N load, friction does not change much. However, as the load increases more, friction increases almost linearly. 3.5 Bearing Inner Friction Torque in Combined Loading. In order to observe the effect of bearing inner friction torque under combined loading, a tilting moment is applied to the bearing. Figure 12 shows the dependence of the friction torque on the tilting moment at 800 N axial loading and at different rotational speeds. The figure presents that as the tilting moment increases, the 021503-4 / Vol. 139, MARCH 2017 Transactions of the ASME Downloaded From: http://tribology.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jotre9/935662/ on 01/31/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use friction torque increases as well. The increase, however, is not linear. In addition, increase in rotational speed causes friction torque to increase without changing the friction torque profile. Dependence of the friction torque on the tilting moment at different axial loads is presented in Fig. 13. At the same rotational speed, increase in tilting moment increases the friction torque. Moreover, increase in axial load increases friction torque." + ] + }, + { + "image_filename": "designv11_34_0003932_s40962-020-00549-5-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003932_s40962-020-00549-5-Figure3-1.png", + "caption": "Figure 3. The 3D model of turbo-expander wheel.", + "texts": [ + " In the present work, we have designed the optimal profile of blade grooves using finite element analysis and established the process of manufacturing the wheel by the plaster casting process using FDM (Figure 2). As shown in Figure 2, we make the plaster mold; its material is PLA, using FDM machine. And then, we manufacture the plaster cores by the plaster mold setup and assemble the plaster cores in the casting mold. Lastly, molten aluminum alloy is poured into the casting mold and the turbo-expander wheel is finished through the machine process. The wheel\u2019s 3D model which was designed with SolidWorks is shown in Figure 3. International Journal of Metalcasting There are 20 blade grooves in the wheel and seal thresholds for air tightness. And in the center there is a hole connected to the rotating shaft. When made of aluminum alloy (A387.0), it weighs 4.4 * 4.6 kg. To improve the efficiency of the wheel, the temperature difference between the air which flows into the input (- 145 to - 149 C) and the air which flows out of the output (- 175 to - 186 C) must be maximum. So the streamline profile of the output must be optimized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002529_10589759.2016.1254211-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002529_10589759.2016.1254211-Figure1-1.png", + "caption": "Figure 1.\u00a0test rig layout.", + "texts": [ + "[37,38] In the present study, the aim is to identify factors governing the generation of vibration signals within the gearbox, particularly as the operational variables such as load and speed change. This investigation focuses mainly on detecting faulty and healthy signals through the use of wavelet packets and multi-layer perceptron and presents the application of the entropy function in feature extraction. The results are applied to the fusion network, which consists of four main multi-layer perceptron networks used to classify healthy or faulty signals in various loads and speeds. The test rig considered in this study is shown in Figure 1: Helical gears have a transmission ratio of 3 and the teeth numbers of 25 and 75. The gears\u2019 pressure angle, surface hardness and roughness are 20o, 600 HV and (2\u20133) \u03bcm, respectively. Table 1 contains the detailed specifications of the gears rotated by three-phase 0.37\u00a0kW asynchronous motors at the nominal speed of 2800\u00a0rpm. To record the vibration signals, two accelerometer sensors of type Br\u00fcel & Kjar (B&K) Model 4507-B-001, 10\u00a0mV/g (1.02\u00a0mV/ (m/s\u00b2)) have been employed. The sensors have a resonance frequency of 513" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000567_ecce.2015.7310281-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000567_ecce.2015.7310281-Figure2-1.png", + "caption": "Fig. 2. Structure transformatio", + "texts": [ + " Finally, section V focuses on the application of this method to a given motor used for railway transportation and we validate the proposed model by comparing the obtained results with the ones coming from both numerical simulations and experimental measurements. 978-1-4673-7151-3/15/$31.00 \u00a92015 IEEE 4397 II. MODELING WITHOUT SATU In this model, the 3D-effects are neglecte considered as composed by two concentric re corresponding to the corrected air-gap region and the Region II corresponding to the P between them. The Maxwell\u2019s equations are s considering a magneto-static problem (in cyli system). We assume that the PM is noncondu Fig. 4 shows a diagram presenting the firstly we transform the real structure i (Slotless) one as shown in Fig. 2; secondly magnetic potential vectors created by the p and stator current (armature reaction), and t superposition principle, we deduce the r potential vector at load-operation; thirdly, we method to account for saturation effect. 1) Simplified representation of the machi To account for the effect of stator op analytical model, the effective air-gap length Carter's coefficient. To illustrate the philosop calculation method is used below [14]. Assu permeability is infinite, and by applying A closed path delimited by Magnetic field line density in the air-gap: where is the air-gap magnetic flux den residual magnetic flux density of the PM at 20\u00b0C, length and PM thickness, and, is the rel of the PM. From (1), we can define the effective air-g The effective air-gap length is then modi coefficient: where is the internal stator radius corre coefficient. Once slot effect is represe coefficient, we describe the spatial distributi mean of Fourier Transform. RATION d. The machine is gions: the region I defined in Fig. 2 M and the space olved analytically ndrical coordinate ctive. analytical model: nto a simplified we determine the ermanent magnets hen, by using the esulting magnetic apply an original ne enings in the 2D is modified with hy behind, the 1D ming that the iron mpere\u2019s law for s, we obtain flux (1) sity, is the are air-gap ative permeability ap length as: (2) fied with Carter\u2019s (3) (4) cted with Carter\u2019s nted by Carter's on of the coils by (5) (6) (7) where is the spatial harmonic : number of the phas are defined in sion structure of the SMPMM n of the electrical machine [15] oils and it's Fourrier Transform for a and short-pitch winding of 5/6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002902_j.cirpj.2020.01.001-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002902_j.cirpj.2020.01.001-Figure1-1.png", + "caption": "Fig. 1. Calculation area digitization scheme.", + "texts": [ + " The process of heat transfer in the deposited layer can be described by means of the non-stationary energy equation: r T\u00f0 \u00de @H @t \u00bc @ @x l T\u00f0 \u00de c\u00f0T\u00de @H @x \u00fe @ @y l T\u00f0 \u00de c\u00f0T\u00de @H @y \u00fe @ @z l T\u00f0 \u00de c\u00f0T\u00de @H @z \u00fe Q\u202f \u00f01\u00de where H is enthalpy, J/kg; t is time, s; r(T) is density of the material, kg/m3; Q is amount of heat released by internal sources per unit volume per unit time when the electron beam acts on the material, W/m3; x, y, and z are Cartesian coordinates of the current calculation point, m; l(T) is thermal conductivity, W/(m K); and c (T) is the specific heat of the material, J/(kg K). To solve the problem, the fully implicit scheme and finitedifference method [18] were used. This scheme leads to a \u201ctridiagonal\u201d system of linear equations that can be solved efficiently by means of the Thomas algorithm\u2014also known as TDMA [19]. The calculation area digitization scheme is shown in Fig. 1. To determine the enthalpy in each reference volume, sequential sweeps along three coordinate axes were repeated for all rows, columns, anddepthsof the temperature array. Each includesaforward and a backward sweep. For this, we used the fractional step method introduced by Yanenko [20]. According to this method, the original multidimensional equation is split into a series of one-dimensional equations that are thensequentiallysolved by TDMA.For example, the first sweep sequence is carried out in the x direction to solve the equation for all y and z nodes (2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001626_j.proeng.2015.12.125-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001626_j.proeng.2015.12.125-Figure1-1.png", + "caption": "Fig. 1. The scheme forming the tooth profile in the face section wheels.", + "texts": [ + " Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/). Peer-review under responsibility of the organizing committee of the International Conference on Industrial Engineering (ICIE-2015) generating circuit [1]. The developed program allows for the processing of cylindrical gears with different parameters of the original contour ( , ha *, hf *, c *, *) and different coefficients of x bias generating circuit rack (Fig. 1). For the further construction of three-dimensional model using a program Autodesk Inventor, with which obtained on the basis of the profile, you can create models of various gears. To create the model spur gear (Fig. 2) is sufficient to use the extrusion operation. In contrast to the spur gear created by standard means of Autodesk Inventor (via the Design Wizard gears) profile tooth surfaces completely fit the profile, really get at run-rack instrument to the settings of the original path. In the case of a helical gear (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002051_j.procir.2016.02.003-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002051_j.procir.2016.02.003-Figure1-1.png", + "caption": "Fig. 1 Structure of a roller bearing Fig. 2: Schematic representation of bearing clearance.", + "texts": [ + " Within this 2D radial cut dimensional as well as geometrical deviations are considered. In section 2 some background information about roller bearings and their clearance are presented. Publications dealing with bearing clearance, the evaluation of bearing clearance or relevant approaches that could be used for this task are discussed in section 3. Afterwards the concept for the evaluation of the bearing clearance is introduced in section 4 and applied in section 5. The paper closes with a conclusion in section 6. 2. Definition of bearing clearance Fig. 1 shows the principal bearing structure for a roller bearing. Although there is a wide variety of types, designs and sizes of roller bearings, they all have in common that they consist of an outer and inner ring. Each of these bearing rings provides a raceway in which the rolling elements roll. For roller bearings the rolling elements are cylinders of slightly higher length than diameter (rollers). The rollers can be guided by a cage or the bearing is designed as a full complement bearing. In this paper only roller bearings of type NU (cf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.4-1.png", + "caption": "Figure 10.4 A 3RRR planar parallel manipulator.", + "texts": [ + "13) If the working space of the parallel manipulator is two dimensional, then \ud835\udf0c consists of the two coordinates of the tip point and the orientation angle of the end-effector. That is, \ud835\udf0c = \u23a1 \u23a2\u23a2\u23a3 p1 p2 \ud835\udf19 \u23a4 \u23a5\u23a5\u23a6 (10.14) A parallel manipulator is called regular if it does not have any kinematic or actuator redundancy or deficiency. For a regular manipulator, it happens that m = \ud835\udf07 = \ud835\udf06. Example 10.4 Position Equations of a 3RRR Planar Parallel Manipulator Consider the 3RRR planar parallel manipulator shown in Figure 10.4. The joint centers A1 and A2 are coincident. The origin O of the base frame is located at the joint center B1. The links of the manipulator are numbered so that 0 is the fixed platform (B1B2B3), 1 and 4 form the first leg L1, 2 and 5 form the second leg L2, 3 and 6 form the third leg L3, and 7 is the moving platform (A1A2A3). The end-effector is a tool shown as QP. The active joints of the manipulator are the three revolute joints located at the points B1, B2, and B3. The constant geometric parameters of the manipulator are listed below", + " b2 = B1B2, b3 = B1B3, b\u2032 3 = B2B3 d1 = A1Q, d3 = A3Q, d7 = A1A3, h7 = QP r1 = B1C1, r2 = B2C2, r3 = B3C3 r4 = C1A1, r5 = C2A2, r6 = C3A3 According to Eqs. (10.6)\u2013(10.9), \ud835\udf07 = 3nm \u2212 2j1 = 3 \u00d7 7 \u2212 2 \u00d7 9 = 3, nikl = j1 \u2212 nm = 9 \u2212 7 = 2 npv = \ud835\udf07 = 3, nsv = 3nikl = 6 The three primary variables are taken as the following active joint variables. \ud835\udf031 = \ud835\udf0301, \ud835\udf032 = \ud835\udf0302, \ud835\udf033 = \ud835\udf0303 Then, the passive joint variables become the six secondary variables. They are expressed as follows by using the orientation angles of the links that are shown in Figure 10.4 and by noting that \ud835\udf19 = \ud835\udf037: \ud835\udf0314 = \ud835\udf034 \u2212 \ud835\udf031, \ud835\udf0325 = \ud835\udf035 \u2212 \ud835\udf032, \ud835\udf0336 = \ud835\udf036 \u2212 \ud835\udf033 \ud835\udf0347 = \ud835\udf19 \u2212 \ud835\udf034, \ud835\udf0357 = \ud835\udf19 \u2212 \ud835\udf035, \ud835\udf0367 = \ud835\udf19 \u2212 \ud835\udf036 The two IKLs can be taken as IKL-1 = B1B2C2A2A1C1B1 and IKL-2 = B1B3C3A3 A2A1C1B1. For these loops, the loop equations can be written as explained below. \u2217 Equation for the link orientations around IKL-1: \ud835\udf0301 + \ud835\udf0314 + \ud835\udf0347 = \ud835\udf0302 + \ud835\udf0325 + \ud835\udf0357 \u21d2 \ud835\udf031 + (\ud835\udf034 \u2212 \ud835\udf031) + (\ud835\udf19 \u2212 \ud835\udf034) = \ud835\udf032 + (\ud835\udf035 \u2212 \ud835\udf032) + (\ud835\udf19 \u2212 \ud835\udf035) \u21d2 0 = 0 (10.15) \u2217 Equation for the link orientations around IKL-2: \ud835\udf0301 + \ud835\udf0314 + \ud835\udf0347 = \ud835\udf0303 + \ud835\udf0336 + \ud835\udf0367 \u21d2 \ud835\udf031 + (\ud835\udf034 \u2212 \ud835\udf031) + (\ud835\udf19 \u2212 \ud835\udf034) = \ud835\udf033 + (\ud835\udf036 \u2212 \ud835\udf033) + (\ud835\udf19 \u2212 \ud835\udf036) \u21d2 0 = 0 (10", + " Therefore, it must be prevented from getting into a pose of PSFK. This can be achieved in two ways. One way is to design the manipulator so that its geometric parameters do not allow any PSFK to occur. The other way is to use the manipulator carefully so that it does not get into a pose of PSFK. In other words, the planned task must be such that it does not give any chance for the primary variables to become interrelated as in Eq. (10.31). Example 10.5 Forward Kinematics of the 3RRR Planar Parallel Manipulator For the 3RRR parallel manipulator shown in Figure 10.4, the scalar loop equations, i.e. Eqs. (10.18)\u2013(10.21), can be written again as shown below. r5c\ud835\udf035 = r4c\ud835\udf034 \u2212 (b2 + r2c\ud835\udf032 \u2212 r1c\ud835\udf031) (10.37) r5s\ud835\udf035 = r4s\ud835\udf034 \u2212 (r2s\ud835\udf032 \u2212 r1s\ud835\udf031) (10.38) d7c\ud835\udf19 = r6c\ud835\udf036 \u2212 (r1c\ud835\udf031 + r4c\ud835\udf034 \u2212 r3c\ud835\udf033 \u2212 b3) (10.39) d7s\ud835\udf19 = r6s\ud835\udf036 \u2212 (r1s\ud835\udf031 + r4s\ud835\udf034 \u2212 r3s\ud835\udf033) (10.40) Position and Velocity Analyses of Parallel Manipulators 353 (a) Forward Kinematic Solution Here, the primary variables (i.e. the angles \ud835\udf031, \ud835\udf032, and \ud835\udf033) are specified and the consequent secondary variables together with the tip point coordinates are found as explained below", + " These positions are illustrated in Figure 10.5 associated with Part (b) about the posture modes. (b) Posture Modes of Forward Kinematics and Posture Mode Changing Poses As seen in Part (a), when actuated by means of the primary angles \ud835\udf031, \ud835\udf032, and \ud835\udf033, the manipulator may be operated in one of the four different posture modes, which are represented by the sign variables \ud835\udf0e1 and \ud835\udf0e2. The relationships between the posture modes and the sign variables \ud835\udf0e1 and \ud835\udf0e2 are explained below. Considering \ud835\udf0e1 with its role in Eq. (10.47) and referring to Figure 10.4 by recalling that \ud835\udf1312 = \u2222( \u2212\u2212\u2192C1C2) and noting that \ud835\udefe12 \u2265 0 (because g12 \u2265 0),\ud835\udf0e1 leads to the following posture modes. They depend on x12 because x12 > 0 makes \ud835\udf1312 an acute angle, whereas x12 < 0 makes \ud835\udf1312 an obtuse angle. They are illustrated schematically in Figure 10.5. PM-1 (A1 above C1C2) \u21d2 \ud835\udf0e1 = +sgn(x12) (10.66) PM-2 (A1 below C1C2) \u21d2 \ud835\udf0e1 = \u2212sgn(x12) (10.67) The borderline between the posture modes considered above is the PMCP (PMCP-12) between them. Such a pose occurs if \ud835\udefe12 = 0 so that \ud835\udf034 = \ud835\udf1312 (10.68) Equation (10.68) implies that the lines C1A1 and C1C2 are aligned. In other words, in PMCP-12, the coincident points A1 and A2 become located right on the line between the points C1 and C2. Note that, in such a critical pose, the manipulator can easily be perturbed to make it continue its subsequent operation in any one of PM-1 and PM-2. Considering \ud835\udf0e2 with its role in Eq. (10.57) and referring to Figure 10.4 by recalling that \ud835\udf1367 = \u2222( \u2212\u2212\u2192C3A1) and noting that \ud835\udefe67 \u2265 0 (because g67 \u2265 0), \ud835\udf0e2 leads to the following posture modes. They depend on x67 because x67 > 0 makes \ud835\udf1367 an acute angle, whereas x67 < 0 makes \ud835\udf1367 an obtuse angle. They are illustrated schematically in Figure 10.5. PM-3 (A3 above C3A1) \u21d2 \ud835\udf0e2 = \u2212sgn(x67) (10.69) PM-4 (A3 below C3A1) \u21d2 \ud835\udf0e2 = +sgn(x67) (10.70) The borderline between the posture modes considered above is the PMCP (PMCP-34) between them. Such a pose occurs if \ud835\udefe67 = 0 so that \ud835\udf036 = \ud835\udf1367 (10", + " In other words, the reachable range of the end-effector and the dexterity of the manipulator can be increased by allowing the PSIKs of the legs to occur. Moreover, when they are allowed to occur due to such tasks, they may be used even in some advantageous ways. For example, if PSIK-k occurs, then Lk becomes useless. So, instead of spending power in order to keep it in an arbitrarily selected pose, it may be allowed to assume a pose caused by its own weight so that no power is spent against gravity. Example 10.6 Inverse Kinematics of the 3RRR Planar Parallel Manipulator For the parallel manipulator shown in Figure 10.4, the following equations can be written in order to express the position of the end-effector by going through each of the three legs. 362 Kinematics of General Spatial Mechanical Systems \u2217 Vector and scalar equations written for the leg L1: \u2212\u2192OP = \u2212\u2212\u2192B1P = \u2212\u2212\u2192B1C1 + \u2212\u2212\u2192C1A1 + \u2212\u2212\u2192A1Q + \u2212\u2192QP \u21d2 p\u20d7 = p1u\u20d71 + p2u\u20d72 = r1u\u20d7(\ud835\udf031) + r4u\u20d7(\ud835\udf034) + d1u\u20d7(\ud835\udf19) + h7u\u20d7 ( \ud835\udf19 + \ud835\udf0b 2 ) (10.112) p1 = r1c\ud835\udf031 + r4c\ud835\udf034 + d1c\ud835\udf19 \u2212 h7s\ud835\udf19 (10.113) p2 = r1s\ud835\udf031 + r4s\ud835\udf034 + d1s\ud835\udf19 + h7c\ud835\udf19 (10.114) \u2217 Vector and scalar equations written for the leg L2: \u2212\u2192OP = \u2212\u2212\u2192B1P = \u2212\u2212\u2192B1B2 + \u2212\u2212\u2192B2C2 + \u2212\u2212\u2192C2A2 + \u2212\u2212\u2192A2Q + \u2212\u2192QP \u21d2 p\u20d7 = p1u\u20d71 + p2u\u20d72 = b2u\u20d71 + r2u\u20d7(\ud835\udf032) + r5u\u20d7(\ud835\udf035) + d1u\u20d7(\ud835\udf19) + h7u\u20d7 ( \ud835\udf19 + \ud835\udf0b 2 ) (10", + " It 364 Kinematics of General Spatial Mechanical Systems is summarized below. \ud835\udf033 = \ud835\udf133 + \ud835\udf0e3\ud835\udefe3 (10.147) \ud835\udf133 = atan2(y3, x3) (10.148) \ud835\udefe3 = atan2(g3, f3) (10.149) f3 = (x2 3 + y2 3 + r2 3 \u2212 r2 6)\u2215(2r3) (10.150) g3 = \u221a x2 3 + y2 3 \u2212 f 2 3 (10.151) \ud835\udf0e3 = \u00b11 (10.152) \ud835\udf036 = atan2[(y3 \u2212 r3s\ud835\udf033), (x3 \u2212 r3c\ud835\udf033)] (10.153) (b) Posture Modes of Inverse Kinematics and Posture Mode Changing Poses \u2217 PML-11, PML-12, and PMCPL-1 of the leg L1: Equation (10.126) implies that \ud835\udf131 = \u2222( \u2212\u2212\u2192B1A1) (10.154) Based on \ud835\udf131, Eq. (10.125) and Figure 10.4 lead to the following conclusions. (i) PML-11 occurs if \ud835\udf031 >\ud835\udf131 > 0 or \ud835\udf031 <\ud835\udf131 < 0. This PML may be called knee-behind. For this PML, \ud835\udf0e1 = +sgn(\ud835\udf131). (ii) PML-12 occurs if \ud835\udf031 <\ud835\udf131 > 0 or \ud835\udf031 >\ud835\udf131 < 0. This PML may be called knee-ahead. For this PML, \ud835\udf0e1 = \u2212sgn(\ud835\udf131). (iii) PMCPL-1 occurs if \ud835\udf031 = \ud835\udf131 or \ud835\udf031 = \ud835\udf131 \u00b1\ud835\udf0b. \u2217 PML-21, PML-22, and PMCPL-2 of the leg L2: Equation (10.137) implies that \ud835\udf132 = \u2222( \u2212\u2212\u2192B2A2) (10.155) Based on \ud835\udf132, Eq. (10.136) and Figure 10.4 lead to the following conclusions. (i) PML-21 occurs if \ud835\udf032 >\ud835\udf132 > 0 or \ud835\udf032 <\ud835\udf132 < 0. This PML may be called knee-behind. For this PML, \ud835\udf0e2 = +sgn(\ud835\udf132). (ii) PML-22 occurs if \ud835\udf032 <\ud835\udf132 > 0 or \ud835\udf032 >\ud835\udf132 < 0. This PML may be called knee-ahead. For this PML, \ud835\udf0e2 = \u2212sgn(\ud835\udf132). (iii) PMCPL-2 occurs if \ud835\udf032 = \ud835\udf132 or \ud835\udf032 = \ud835\udf132 \u00b1\ud835\udf0b. \u2217 PML-31, PML-32, and PMCPL-3 of the leg L3: Equation (10.148) implies that \ud835\udf133 = \u2222( \u2212\u2212\u2192B3A3) (10.156) Based on \ud835\udf133, Eq. (10.147) and Figure 10.4 lead to the following conclusions. (i) PML-31 occurs if \ud835\udf033 >\ud835\udf133 > 0 or \ud835\udf033 <\ud835\udf133 < 0. This PML may be called knee-behind. For this PML, \ud835\udf0e3 = +sgn(\ud835\udf133). (ii) PML-32 occurs if \ud835\udf033 <\ud835\udf133 > 0 or \ud835\udf033 >\ud835\udf133 < 0. This PML may be called knee-ahead. For this PML, \ud835\udf0e3 = \u2212 sgn(\ud835\udf133). (iii) PMCPL-3 occurs if \ud835\udf033 = \ud835\udf133 or \ud835\udf033 = \ud835\udf133 \u00b1\ud835\udf0b. The posture modes and thePMCPs of the legs are illustrated in Figures 10.8 and 10.9. (c) Position Singularities of Inverse Kinematics As mentioned before, each leg may have a PSIK of its own. The PSIK of Lk is denoted here as PSIK-k", + " (10.179), they are obtained as follows by differentiating Eq. (10.11) G\u0302x(x, y)x\u0307 + G\u0302y(x, y)y\u0307 = 0 (10.195) In Eq. (10.195), G\u0302x(x, y) and G\u0302y(x, y) are defined so that [G\u0302x(x, y)]ij = Gxij(x, y) = \ud835\udf15gi(x, y)\u2215\ud835\udf15xj (10.196) [G\u0302y(x, y)]ij = Gyij(x, y) = \ud835\udf15gi(x, y)\u2215\ud835\udf15yj (10.197) Upon comparing Eqs. (10.195) and (10.179), it is seen that M\u0302(x, y) = G\u0302y(x, y) (10.198) N\u0302(x, y) = \u2212G\u0302x(x, y) (10.199) Example 10.7 Velocity Equations of the 3RRR Planar Parallel Manipulator Consider the manipulator shown in Figure 10.4. For this manipulator, the position relationships have already been expressed by Eqs. (10.27) and (10.28). These equations lead to the following velocity equations upon differentiation. Position and Velocity Analyses of Parallel Manipulators 371 Differentiation of Eq. (10.27): [ v1 v2 ] = [ p\u03071 p\u03072 ] = [ \u2212(r1s\ud835\udf031)?\u0307?1 \u2212 (r4s\ud835\udf034)?\u0307?4 \u2212 (d1s\ud835\udf19 + h7c\ud835\udf19)?\u0307? +(r1c\ud835\udf031)?\u0307?1 + (r4c\ud835\udf034)?\u0307?4 + (d1c\ud835\udf19 \u2212 h7s\ud835\udf19)?\u0307? ] Velocity equation for the tip point: [ v1 v2 ] = [ \u2212r1s\ud835\udf031 +r1c\ud835\udf031 ] ?\u0307?1 + [ \u2212r4s\ud835\udf034 +r4c\ud835\udf034 ] ?\u0307?4 \u2212 [ h7c\ud835\udf19 + d1s\ud835\udf19 h7s\ud835\udf19 \u2212 d1c\ud835\udf19 ] ", + " Besides, at the instant of MSIK-k, the actuator or actuators of Lk can be used in order to select the posture mode (PML) of Lk after the MSIK-k. This is because the pose of MSIK-k happens to be the same as the PMCPL of Lk , i.e. PMCPL-k. Thus, with a little actuation effort, the current PML of Lk can either be maintained or it can be changed into another PML for the post-singularity motion of the manipulator. 380 Kinematics of General Spatial Mechanical Systems Example 10.9 Inverse Velocity Analysis of the 3RRR Planar Parallel Manipulator For the parallel manipulator shown in Figure 10.4, the following velocity equations can be obtained by differentiating the position equations written for each leg in Example 10.6. \u2217 Velocity equations for L1: v1 = \u2212(r1s\ud835\udf031)?\u0307?1 \u2212 (r4s\ud835\udf034)?\u0307?4 \u2212 (d1s\ud835\udf19 + h7c\ud835\udf19)?\u0307? (10.265) v2 = +(r1c\ud835\udf031)?\u0307?1 + (r4c\ud835\udf034)?\u0307?4 + (d1c\ud835\udf19 \u2212 h7s\ud835\udf19)?\u0307? (10.266) \u2217 Velocity equations for L2: v1 = \u2212(r2s\ud835\udf032)?\u0307?2 \u2212 (r5s\ud835\udf035)?\u0307?5 \u2212 (d1s\ud835\udf19 + h7c\ud835\udf19)?\u0307? (10.267) v2 = +(r2c\ud835\udf032)?\u0307?2 + (r5c\ud835\udf035)?\u0307?5 + (d1c\ud835\udf19 \u2212 h7s\ud835\udf19)?\u0307? (10.268) \u2217 Velocity equations for L3: v1 = \u2212(r3s\ud835\udf033)?\u0307?3 \u2212 (r6s\ud835\udf036)?\u0307?6 + (d3s\ud835\udf19 \u2212 h7c\ud835\udf19)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003108_1350650120925363-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003108_1350650120925363-Figure2-1.png", + "caption": "Figure 2. Proceed novel GFCB and its application on the air compressor.", + "texts": [ + " Bearing structures and test rigs The structure of the novel GFCB is shown in Figure 1. The main difference between the new and the traditional GFB is that the inner wall of the housing is made of conical type, which makes the bearing with large end and small end can offer the radial and axial load capacity at the same time. The bump foil acts as elastic support, and the top foil provides the lubricating gas film. Moreover, the friction between foils and sleeve can provide some friction damping so that the GFCB will absorb some vibration in the running. Figure 2(a) shows the processed GFCB and the shaft parts matched with it, in which a chute on the bearing sleeve has been processed to fix the top foil and bump foil. The shaft parts include the front cone segment, the rear cone segment, magnetic steel, and intermediate protective sleeve. Figure 2(a) also shows the mold used for making bump foil, in which a piece of polytetrafluoroethylene (PTEE) is used for the stamping forming by the mold. To verify the practical application feasibility of the novel bearing proposed in this article, a pair of GFCBs was used to replace the traditional radial-thrust foil bearing supporting scheme of an air compressor, as shown in Figure 2(b). And two vibration sensors were placed at the position of front and rear conical bearings respectively to monitor the vibration of the air compressor during the actual operation, which could be used to judge the actual supporting effect of GFCB. In the conical coordinate system shown in Figure 3, the Reynolds equation for the novel GFCB is @ @ h3 @p2 @ \u00fe r sin @ @r r sin h3 @p2 @r \u00bc 12u!r2 sin2 @ ph\u00f0 \u00de @ \u00fe 24ur2 sin2 @ ph\u00f0 \u00de @t \u00f01\u00de where and r are circumferential and radial coordinate, respectively, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000400_978-0-8176-4893-0_7-Figure7.18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000400_978-0-8176-4893-0_7-Figure7.18-1.png", + "caption": "Fig. 7.18 The pendulum-cart system", + "texts": [ + " The comparison of effectiveness of HOSM-based uncertainty identification and compensation versus sliding mode based uncertainty compensation, is presented in [79]. HOSM observer based control for the compensation of unmatched uncertainties was developed in [80]. State estimation and input reconstruction in nonminimum phase causal nonlinear systems using higher-order sliding mode observers is studied in [166]. Automotive applications of sliding mode disturbance observer- based control can be found in the book [111]. Exercise 7.1. The pendulum-cart system (see Fig. 7.18), when restricted to a twodimensional motion, can be described by the following set of equations: .M Cm/ Px Cml R D u.I Cml2/ R Cml Rx D mgl (7.74) where M and m are the mass of the cart and the pendulum, respectively, l is the pendulum length, .t/ is its deviation from the vertical, and x.t/ represents the horizontal displacement of the cart. The system parameters are given asM D 2\u0152kg ; m D 0:1\u0152kg , and l D 0:5\u0152m . Given the measured outputs and x, design a supertwisting observer for P and Px for the uncontrolled case, u D 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003589_1350650120962973-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003589_1350650120962973-Figure2-1.png", + "caption": "Figure 2. The equivalent normal deviation e\u00f0Mi\u00de at point Mi.", + "texts": [ + "32 The deflection function at any potential point of contact Mi on the tooth flanks can be expressed as follows:33,34 D\u00f0Mi\u00de \u00bc \u00bdVG\u00f0Mi\u00de T\u00bdX e\u00f0Mi\u00de (1) where X \u00bc \u00bdv1;w1; h1; vg;wg; hg; vp;wp; hp; v2;w2; h2 T represents the DOFs can be ex VG\u00f0Mi\u00de is the pinion-gear structural vector associated with point Mi which contains the pinion-gear geometrical properties at Mi (see equation (2)); e\u00f0Mi\u00de is the equivalent normal deviation at point Mi with respect to ideal flanks (theoretical contact lines) and accounts for geometrical errors (see Figure 2). \u00bdVG\u00f0Mi\u00de \u00bc \u00bd0; 0; 0; sin\u00f0a\u00de; cos\u00f0a\u00de;Rb1; sin\u00f0a\u00de; cos\u00f0a\u00de;Rb2; 0; 0; 0 T (2) where a represents the pressure angle, Rb1 and Rb2 are the gear and pinion base radii. The contact condition for each pointMi in the base plane deduced from equation (1) is as follows: \u2022 D\u00f0Mi\u00de > 0 \u2014\u2014> Contact at point Mi \u2022 D\u00f0Mi\u00de < 0 \u2014\u2014> Lost of the contact at point Mi The error profile is specified as the distance separating the real tooth profile and the theoretical tooth profile. Two cases of interest are considered: a constant error profile is introduced in local friction coefficient formulation (see Figure 3) and a variable error profile along the action contact line and the theoretical contact line is considered (see Figure 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003595_icuas48674.2020.9214014-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003595_icuas48674.2020.9214014-Figure2-1.png", + "caption": "Figure 2. Illustration of the hover-to-cruise transition process", + "texts": [ + "com ) Hailong Pei is with Key Laboratory of Autonomous Systems and Networked Control, Ministry of Education, Unmanned Aerial Vehicle Systems Engineering Technology Research Center of Guangdong, College of Automation Science and Engineering, South China University of Technology, Guangzhou, 510640,China (e-mail: auhlpei@scut.edu.cn, corresponding author: +86-13660182029) Since high-speed level flight is one of the key characteristics, it is of great significance to study the hover-to high-speed transition control of a ducted fan UAV (See Fig.2). For near-hover control, plenty of previous works have shown good performances of their design [4]-[6]. However, as for the hover-to-high-speed transition process, highly nonlinearity of the dynamic model and complexity of the components put great challenge on the controller design [7]. All aerodynamic effects suffer great change during the process. On one hand, the equilibrium point of the dynamic system encounters a great deviation from hover to level flight with highly nonlinearity. On the other hand, as the forward speed increasing, the aircraft becomes more and more sensitive to inputs as well as disturbances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001795_robio.2015.7418892-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001795_robio.2015.7418892-Figure2-1.png", + "caption": "Figure 2. The sketch of inverse kinematic with geometry method.", + "texts": [ + " So the homogeneous transformation matrix that relates coordinate system {W } to the base coordinate system can be expressed as 0 1 2 3 4 5 6 7 0 1 2 3 4 5 6 7 G G W WT T T T T T T T T T (1) B. Inverse Kinematic with Geometry Method The manipulator is fixed on base coordinate system and share the same coordinate system with target object. To grasp the object, the manipulator needs to construct its body shape according to the position and orientation of target object. After the shape is definite, the rotational angle of each joint can be solved according to the geometrical relationship [13]. As shown in Fig 2, in base coordinate system, the coordinates of joint 0 and point A are 0 0 0 and 10 0 l respectively. The target object\u2019s coordinates is X Y ZP P P and the target object\u2019s coordinate system is T T T TO X Y Z . To grasp target object successfully, the position and orientation of terminal gripper should overlap with that of target object. So the coordinates of point C in target object coordinate system is 4 0 0l . According to the relationship between target object coordinate system and base coordinate system, the coordinates of point C in base coordinate system can be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000972_amm.711.27-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000972_amm.711.27-Figure1-1.png", + "caption": "Fig. 1 Schematic of a plain journal bearing", + "texts": [ + " Gao [7] and Zhang [8] studied the design of hydrodynamic water-lubricated journal bearing and step thrust bearing, respectively. But these researches all focused on water lubrication, few have compared the lubrication performances of water-lubricated bearing and oil-lubricated bearing quantitatively. In order to show differences of lubrication performances of the two bearings clearly, this paper calculated the flow models of water-lubricated and oil-lubricated plain journal bearings using CFD method. Plain journal bearing model Figure 1 shows a schematic of a plain journal bearing and the coordinate system. The bearing is fully submersed in water or oil and is subjected to a constant external vertical load. There is a radial clearance (c) between the journal and bearing. As the rotating journal deviates from the bearing centre, dynamic pressure is generated in the thin water film to counteract the external load. At the equilibrium position, the distance between the journal centre and the bearing centre is the journal eccentricity (e) and the angle between the eccentricity orientation and the load direction is the attitude angle (\u03d5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000415_3527600906.mcb.20130069-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000415_3527600906.mcb.20130069-Figure11-1.png", + "caption": "Fig. 11 Schematic of a piezoelectric transducer.", + "texts": [ + " For an acoustically thin film, the mass change affects the transducer response [93, 94], whereas for an acoustically thick film, the film\u2019s viscous and elastic properties and geometric features also make significant contributions. Generally, the change in mass is central to the application of piezoelectric transducers in biosensors. However, there are instances when the ability of such transducers to quantify changes in shear modulus and viscosity has been exploited to fabricate efficient biosensors to study lipids and membranes. Figure 11 shows a typical piezoelectric transducer; here, the biosensing layer with the immobilized bioreceptors can be fabricated over the transducer surface. Piezoelectric biosensors have been employed in the label-free detection of a wide array of analytes ranging from proteins, oligonucleotides and DNAs, antigens, small molecules to viruses and bacteria. They have also been used widely to study protein\u2013protein, protein\u2013DNA, protein\u2013peptide, peptide\u2013peptide interactions, as well as interactions of carbohydrates with proteins, lipids, and other carbohydrates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003739_0954405420971128-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003739_0954405420971128-Figure5-1.png", + "caption": "Figure 5. Laboratory press-tool system utilised in coin minting of additive manufactured coin blanks: (a) scheme of the press-tool system with the geometry of the coin blanks and (b) detail and photograph of the dies utilised in coin minting (\u2018Eagle\u2019 and \u2018Rooster\u2019).", + "texts": [ + " The holes were not polished although they could have been by subjecting the built-up cylinders to abrasive flow machining. Figure 4 shows a scheme of the deposition strategy (refer to \u20181\u2019, \u20182\u2019, etc.) together with photographs after SLM, wire-EDM and polishing, in which the improvement of surface finishing is easily observed. Deposition parameters used in the fabrication of coin blanks by SLM were identical to those provided in Table 1. The coin blanks were then compressed (coin minting) between dies in a press-tool (Figure 5) to impart lettering and other reliefs on both surfaces. The tool was designed and fabricated by the authors and consists of a simple laboratory system with an upper drive plate, a lower fixed plate, an upper die holder, a pair of reverse and obverse dies and a collar with an inner flat surface (Figure 5(a)). The dies have a central recess of 0.15mm to avoid contact in these regions of the \u2018Rooster\u2019 and \u2018Eagle\u2019 coins (refer to the central blue area in the reverse and obverse die surfaces of Figure 5(b)). They were made from an AISI D3 cold work steel, also designated as DIN 1.2080, which was heat treated (hardening and tempering) to ensure a hardness of 60 HRc. The collar was made of tungsten carbide. The coin minting experiments were carried out in the same hydraulic testing machine in which material compression tests were performed. In each test, the coin blanks were compressed up to the required end of stroke in order to replicate coin minting by a single strike. At least five repetitions were made for each prototype coin design", + "000 hexahedral elements with four layers of elements across thickness and a higher density of elements in the outer region where lettering was imparted. The dies and collar were modelled as rigid objects and were discretised by means of spatial triangular contact-friction elements. The meshes at centre of the dies are slightly recessed, as in the actual dies, so that contact with the coin blanks preferentially takes place at the outer region where lettering and additional design features are located (Figure 5). Computer implementation of the finite element flow formulation in i-form required parallelisation of the subroutines handling the solution of the non-linear set of equations resulting from the discretisation of (3) by means of hexahedral elements. The parallelisation was carried out in OpenMP (open shared-memory multiprocessing) and details are given in Nielsen et al.17 The central processing unit (CPU) time for a typical analysis using a convergence criteria for the velocity field and residual force equal to 1022 was approximately 48 h on a computer equipped with an Intel i75930K CPU processor with six cores", + " This is done by analysing the finite element predicted evolution of the compressive z-stresses (hereafter designated as the \u2018pressure\u2019) in the reverse and obverse surfaces of coin samples corresponding to the four different percentages of the total die stroke that are disclosed in Figure 8. The dark red colour corresponds to pressures close to zero and, therefore, to regions not in contact with the dies. The other colours evolving from light red to dark blue correspond to increasing values of pressure applied by the dies. As seen, the centre of both coins remains free of contact up to the end of stroke due to the recesses of 0.15mm existing in both die surfaces (Figure 5). In contrast, the outer region plastically deforms to fill the intricate details of lettering and other reliefs, and to adjust the diameter of the coin to the geometry of the collar. At the end of stroke, the finite element predicted geometries of the prototype coins are very close to the actual ones (Figure 9). The second topic to be addressed is related to material flow at the centre of the coin blanks containing the complex and intricate contoured holes. Figure 10 shows a top view of the normalised velocity field v= vxy v0j j where vxy is the resultant xy-velocity of the material, and v0 is the velocity of the upper drive tool plate (upper die)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002995_s10957-020-01652-7-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002995_s10957-020-01652-7-Figure13-1.png", + "caption": "Fig. 13 The inverted pendulum system", + "texts": [ + " Figure 11 compares the value of the cost function V0, V1 and V \u2217 based on the suboptimal control policy in this paper, the control policy in [43] (linear quadratic regulator with optimal output feedback solution algorithm) and the optimal control policy, respectively. Figure 12 shows the convergence of the policy iteration. From Fig. 12, it can be found that the proposed PI algorithm converges within four iterations and can be achieved smaller control input u = 32.885. As a comparison, the algorithm in [43] gives a control input u = 33.728. Plainly, by using the proposed PI algorithm, the control input energy is reduced 2.499% Example 3 Consider the inverted pendulum system is shown in Fig. 13. The dynamic behavior can be described as follows [44,45]: x\u03071(t) = x2(t) x\u03072(t) = f1(x, t) + b1(x, t)u1(t) + d1(t) x\u03073(t) = x4(t) x\u03074(t) = f2(x, t) + b2(x, t)u2(t) + d2(t), (64) where f1(x, t) = mt g sin(x1)\u2212mpL sin(x1) cos(x1)x22 L(4mt/3\u2212mpcos2(x1)) , b1(x, t) = cos(x1) L(4mt/3\u2212mpcos2(x1)) , f2(x, t) = \u22124mpLx22 sin(x1)/3\u2212mpg sin(x1) cos(x1) 4mt/3\u2212mpcos2(x1) , b2(x, t) = 4 4mt\u22123mpcos2(x1) . The physical meanings and the values of the parameters are shown in Table. 1. The disturbances are chosen as d1(t) = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure9-1.png", + "caption": "Fig. 9 RL1 Hand\u2019s driving system", + "texts": [ + " It requires not only moving the hand to the necessary location but also a previous step to put the hand in an operative state. For that reason, the RL1 Hand should pass through an intermediate step that allows it to leave the anchor in order to be operative afterwards. The actuator system should provide all the necessary states the RL1 Hand might need to achieve holding tasks and, apart from that, should develop all the movements to pass from the different states: standby to operative. A specific design for the RL1 Hand of an actuator system is shown in Fig. 9. The driving system of the RL1 Hand is formed by an energy transformer that provides three states to the two mechanisms named MA1 and MA2. Both mechanisms give movement to all the robotic system elements. Only one of them works at a time, thanks to a specific mechanism. This allows the energy transformer to send its movement selectively to each Max depending on the state of the Hand. The extraction system is shown in Fig. 10. In it is possible to see, apart from all the elements that form it, a mobile platform named Main Platform, which moves in only one direction through the Guiding Columns", + " Gonz\u00e1lez Rodr\u00edguez Legs are the most important and active member among the ones that improve the mobility of robots. Unlike wheeled robots they must be able to move in soft and rough terrains (like mud, sand, forest, rocks) or overcoming obstacles like steps. With regard to the configurations of legs for walking robots, there are three types of research lines: articulated legs, passive-dynamic legs, and decoupled configuration legs. The vast majority of the current legged robots use very simple leg schemes, which are made up of a chain of links joined by joints according to one of the two forms shown in Fig. 9: insect-like legs and mammal-like legs. The legs of reptiles have a similar configuration to mammals, although the parallel axes (labelled as T2 and T3) belong to very separated vertical planes and hence they undergo higher torques (Guardabrazo and Gonzalez de Santos 2004). Reptile-like robots also require an articulated body to keep balance during the crawl movement (Ishihara and Kuroi 2006). Although different types of actuators can be found, most of biped robots use the scheme shown in Fig. 9b in which electrical motors drive rotational joints. As reproduced in Fig. 10a biped robot legs require a relatively wide and flat foot to provide static stability and improve the dynamic one. Unless a penguin gait is not a concern (Collins et al. 2005), Mobile Robots 49 the clearance of the foot must be accomplished by means of an ankle, with a rotational joint whose axis is perpendicular to the sagittal plane. In order to keep balance a lateral movement is required that is performed by a set of two joints per leg (not shown in the figure), one in the hip and another in the ankle, whose axes are parallel to the movement direction", + " 2006) since for the design of an anthropomorphic hand a suitable knowledge of the human grasp is needed. Cylindrical grasp of human hand has been analysed since it represents the most used grasp in industrial applications (Cutkosky 1989). In particular, dimensions of fingers, grasping forces and contact points between fingers and objects have been investigated in human grasping. The dimensions of each phalange of index, medium and thumb have been measured for five persons (Nava Rodr\u00edguez et al. 2004). Figure 9a shows a photo sequence of a cylindrical grasp performance of a human hand. The positions of marks of human hand fingers of Fig. 9a have been measured in order to achieve this position by the robotic hand of Fig. 8. The force measuring system of Fig. 9b has been set up in a human hand for developing the experimental test of Fig. 9c. The grasp forces of several objects with different shapes and dimensions have been measured in order to compare results with experimental validations of the robotic hand of Fig. 8. The experimental results show the practical feasibility of the prototype as three-fingered robotic sensored hand with three 1 DOF anthropomorphic fingers, having human-like operation (Nava Rodr\u00edguez et al. 2004). The human-like characteristics of the robotic hand of Fig. 8 simplify its control architecture since some aspects of grasp have been checked in the mechanical design, for example performance of fingers and grasping force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003829_s10846-020-01290-1-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003829_s10846-020-01290-1-Figure17-1.png", + "caption": "Fig. 17 Experiment results of two cylindrical components. a and c in camera 1; b and d in camera 2. The red point represents the movement step. The green and blue points represent the goal positions with and without disturbance, respectively", + "texts": [ + " But the relative movements in mid view are point to the goal position directly. The two relative movements in image spaces are transferred to 3D space with the method in Section 3.6. But the optimized variable in (14) is only \u03b52. The results are given in Figs. 15 and 16. The trajectories are smooth. It shows good adaptability of our method when meeting different kinds of components. Our method can be used not only in needle reaching tasks, but also in reaching tasks with other components. A set of experiments are conducted. In Fig. 17a and b, the results of four experiments are given. The cylindrical component can move to the goal position smoothly from different start positions. In Fig. 17c and d, the cylindrical component is lifted at t = t0. Our method can still work well although the position of component changes suddenly. The trajectories in 3D space are shown in Fig. 18. It shows that our method has wider application fields. The reaching skill learning framework is proposed for components based on imitation learning. An image segmentation module is adopted to distinguish different components. A neighborhood-sampling module is proposed for extracting features effectively and an attention module is adopted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002814_lra.2020.2965914-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002814_lra.2020.2965914-Figure5-1.png", + "caption": "Fig. 5. Top: Hardware components. Bottom: Hand kinematics. Wrist and finger actuation tendons run antagonistically to each other, with the latter propagating its force to the six individual phalanges through a whiffletree differential with segment ratios shown.", + "texts": [ + " IMPLEMENTATION To realize a prototype hand with a bending joint close to the fingers that is still capable of exerting high torques, the finger and wrist actuation are coupled with a tendon network resulting in two controlled degrees of freedom. A three-finger design was chosen to automatically achieve statically determinant grasps in many situations. Two phalanges per finger are sufficient to achieve both the \u201cpower pinch\u201d grasp discussed above and a more traditional enveloping grasp. Proximal and distal phalange lengths were set to 50 mm and 35 mm respectively, resulting in an overall length similar to human fingers and a ratio that worked well with the rim heights of most dishware. Fig. 5 shows the kinematic descriptions as well as the components of the final implementation. With the wrist and finger actuation tendons running antagonistically to each other, it is necessary to regulate their forces such that grasp force is maintained during wrist movement. Two sensors were integrated to provide feedback for this process. A potentiometer is anchored in the wrist joint to provide absolute position information on the angle of the wrist (Fig. 5E), and a strain gauge anchors the finger tendon redirect pulley to provide a force signal (Fig. 5J). These features combine to allow a position value to be specified by the hand controller while the fingers can be actuated with force feedback to achieve a desired grasp strength at any reachable wrist position. Force ratios at each phalange are, in turn, maintained by the length ratios of a floating whiffletree apparatus mounted between the finger actuator and the phalanges (Fig. 5F). For a symmetric grasp force, the first portion of the whiffletree distributes half of the total force to the thumb and one quarter of the total force to each of the opposing fingers. The final layer of the whiffletree allocates force to the distal and proximal phalange anchored tendons. Considering the loading scenario presented in Fig. 4, it is desirable to locate the center of pressure of a pinching finger in the middle of the distal phalange. It is shown in Eqn. 3 that this Authorized licensed use limited to: University of Wollongong", + " In turn, these low motor torques reduced manipulation speed and limited grasp force, which, in some circumstances, prevented the fingers from settling into the full planar contact condition desired. Future prototypes will avoid this difficulty with stronger materials, allowing the motors to be actuated at a higher percentage of their total capacity and thus improve grasp security significantly. Finally, this prototype is designed to be used in conjunction with a sensory skin across all surfaces of the phalanges and palm (note the associated PCB housed in the design in Fig. 5D). These sensors are under construction and will be integrated in the near future for grasp sensing and the detection of accidental contacts. Toyota Research Institute provided funds to assist the authors with their research but this letter solely reflects the opinions and conclusions of its authors and not TRI or any other Toyota entity. REFERENCES [1] R. B. Rusu, B. Gerkey, and M. Beetz, \u201cRobots in the kitchen: Exploiting ubiquitous sensing and actuation,\u201d Robot. Auton. Syst., vol. 56, no. 10, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000422_transducers.2015.7180945-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000422_transducers.2015.7180945-Figure3-1.png", + "caption": "Figure 3: Image of the piezo, lens body with substrate, finished lens, and packaged lens.", + "texts": [ + " Still, we so far only described a membrane. To create a lens, we need to add a refractive medium below the membrane. In [7], this was done with a fluid, but now, we will join the membrane to an elastomer lens body as shown in fig. 2. Because of the counter-pressure upon deformation, the thickness and elasticity of this body will also influence the tuning range. 978-1-4799-8955-3/15/$31.00 \u00a92015 IEEE 399 Transducers 2015, Anchorage, Alaska, USA, June 21-25, 2015 The fabrication is a straightforward pick-and-place process (fig. 3): First the piezo and the glass are structured by laser ablation and glued together using a rigid polyurethane. The elastomer body is molded directly onto the lens substrate in a milled polymer mold. Then, the membrane is glued on top of the body using the same elastomer, which fills the optical defects caused by the relatively rough mold. The lens is finally placed in a compact package (fig. 4) and equipped with soldering contacts. We have fabricated prototypes using 30 \u00b5m thin borosilicate glass and an approximately 120 \u00b5m thick transversely polarized PZT piezo film" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002305_b978-0-12-405935-1.00010-1-Figure10.17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002305_b978-0-12-405935-1.00010-1-Figure10.17-1.png", + "caption": "FIGURE 10.17 Streamline pattern near the separation point (x \u00bc xs, y \u00bc 0) on a flat surface.", + "texts": [ + " Furthermore, ideal-flow theory may not be used to determine the pressure in a separated flow region, since the flow there is rotational and the interface between irrotational and rotational flow regions no longer follows the body\u2019s solid surface. Instead, the irrotationalrotational flow interface may be some unknown shape encompassing part of the body\u2019s contour, the separation streamline, and, possibly, a wake-zone contour. EXAMPLE 10.8 Using a third-order two-dimensional power-series expansion near a flat-plate boundary layer\u2019s separation point, x \u00bc xs and y \u00bc 0, determine how the stream function j(x,y) depends on vp/vx and bs, the angle the separating streamline makes with the horizontal surface as shown in Figure 10.17. Solution A third-order power series expansion for j(x,y) is: j x; y \u00bc a0 \u00fe a1x0 \u00fe a2y\u00fe a3x0 2 \u00fe a4x0y\u00fe Ay2 \u00fe a5x0 3 \u00fe a6x0 2y\u00fe Bx0y2 \u00fe Cy3: where x0 \u00bc x e xs, and a0 through a6, A, B, and C are undetermined constants. This stream function must satisfy the no-slip boundary condition, u \u00bc v \u00bc 0 on y \u00bc 0, so vj/vy \u00bc evj/vx \u00bc 0 on y \u00bc 0. These two conditions cause a1 through a6 to be zero, and if j \u00bc 0 defines the plate surface, then the stream function reduces to j\u00f0x; y\u00de \u00bc Ay2 \u00fe B\u00f0x xs\u00dey2 \u00fe Cy3. In addition, the surface shear stress, sw, is zero at the separation point, so: sw \u00bc m vu vy y\u00bc0;x\u00bcxs \u00bc m v2j vy2 y\u00bc0;x\u00bcxs \u00bc \u00f02A\u00fe 2B\u00f0x xs\u00de \u00fe 6Cy\u00dey\u00bc0;x\u00bcxs \u00bc 2A \u00bc 0; and this leaves: j x; y \u00bc B x xs y2 \u00fe Cy3: In the vicinity of the separation point, this stream function j(x,y) must satisfy two additional conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002539_iecon.2013.6699805-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002539_iecon.2013.6699805-Figure3-1.png", + "caption": "Fig. 3. Coordinate Frames and Hovering-Mode Operational Principles", + "texts": [ + " i) A Commercial Off-The Shelf (COTS) USB camera for estimation of the UAV\u2019s translational motion. j) 3 high-torque Servos, controlling the rotors\u2019 orientation. k) 2 high-power Scorpion SII-v2-3020 780kV Brushless DC (BLDC) main motors with 3-bladed 13x8-inch propellers for the main rotors, and a Robbe Roxxy 2827-34 tail motor with a 3-bladed 9x7-inch propeller. l) 3 custom-programmed fast-response Electronic Speed Controllers (ESCs) controlling the motors\u2019 rotational speeds. The UPAT-TTR\u2019s translational hovering mode operation principles are presented in Figure 3. Also depicted are the Body-Fixed coordinates Frame (BFF) B = {Bx, By, Bz} and the North-East-Down (NED) [3] Local Tangential coordinates Plane (LTP) E = {N, E, D}. Let \u0398 = {\u03c6 , \u03b8 , \u03c8} be the LTP-based rotation angles vector, and XW = {x, y, z} the LTP-based position vector. Also, let \u2126 = {p, q, r} be the BFF-based angular rotation rate vector, and U = {u, v, w} be the BFF-based velocity vector. In achieving the system\u2019s rotational (attitude) control, the roll (\u03c6 ) is controlled via the differential thrusting of the main rotors, the pitch (\u03b8 ) via the differential thrusting of the front and tail rotors, and the yaw (\u03c8) via the tilting of the tail rotor. Also, as depicted in Figure 3, the translational motion is controlled as follows: The vertical (Bz direction) translation is controlled via the total thrust produced by the 3 rotors. The lateral (By direction) translation is controlled via the projection of the total thrust vector in that direction, which is achieved by producing a roll (\u03c6 ) angle. The longitudinal (Bx direction) translation is controlled via tilting the main rotors by an angle \u03b3x , thus projecting/vectoring their thrust along the longitudinal degree-of-freedom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000230_ijcaet.2017.086921-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000230_ijcaet.2017.086921-Figure1-1.png", + "caption": "Figure 1 Suspension system of rail road vehicle", + "texts": [], + "surrounding_texts": [ + "The spring has free play up to its solid height as it can achieve its deformation at static as well as in dynamic way up to this limit. It is observed that the end axle spring may deflected up to 76 mm as it achieve solid height after 76 mm, accordingly other springs deflections are restricted up to its height as end axle spring is 20 mm less in height than middle axle outer spring as shown in Table 1, technical specifications of primary spring. It is considered that springs may deform up to maximum height in dynamic condition at its maximum speed of 80 km/hr during curving, tracking or on uphill. By considering its deformation, the forces on each spring are calculated by considering the rails are smooth and there are no geometrical irregularities. The effects of deformation are acquired by each primary spring by reducing height of whole suspension system due to variation in free heights of every spring as shown in Table 1. As per the deflection of spring (Michalczyk, 2009), forces can be calculated for all springs by using following formula. The forces are also calculated for actual deformation of spring of rail road vehicle in static and dynamic condition, i.e., for the vehicle moving at straight track, curved track and at uphill. The deformations of spring are also verified in shed when rail vehicle is at rest, i.e., at static condition. The possible deflections and forces on each primary spring for reduced height are shown in Table 2 and forces calculated for actual deformation of springs are shown in Table 3. 3 m 4 8FD n\u03b4 Gd = (1)" + ] + }, + { + "image_filename": "designv11_34_0003970_ssrr50563.2020.9292576-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003970_ssrr50563.2020.9292576-Figure2-1.png", + "caption": "Fig. 2. Flexible continuous track without passive rollers", + "texts": [ + " The other common methods to climb high obstacles are increasing the number of drive wheels, adding flippers of a continuous track, and introducing a wheel composed of a variable-diameter mechanism. In these methods, however, more actuators and axles are necessary, increasing the complexity of the configuration of the robot. To overcome these problems, we proposed a novel mobile mechanism, called mono-wheel track (MW-Track), as indicated in Fig. 1. The structure of the MW-Track is unique as it was composed of an elastic track belt driven by a single active wheel, instead of passive rollers, which are usually employed in the general continuous track (Fig. 2). As the track is flexible and not strained, it can be easily deformed by external forces; therefore, it can adapt to the shape of the terrain. This new configuration has the following characteristics: (i) simple structure, as additional actuators or axles, such as flippers, are not required for climbing high obstacles; (ii) high mobility on rough terrain; (iii) efficient transmission of the driving force to the ground, as the adherence to the ground is high owing to several contact points to the ground that prevent it from spinning idly", + " The sprocket guard covers the front of this element to prevent the track belt from touching the sprocket and become jammed. The grousers are 3D printed and covered by rubber on the surface. The grousers are fixed to the chain attachment with screws at regular intervals to avoid contact when the track belt is bent at 90\u00b0. Because the angle of the grouser easily changes, owing to the deformation of the crawler, the shape of the grouser was rounded, resulting in a constant frictional force. Optimal grouser shapes that can be hooked to various surfaces will be examined in the future. The experimental robot shown in Fig. 2 is composed of two units of the MW-Track, two motors, a tail, and a body frame. The tail and the body frame are parallel to the track holder; therefore, the robot is vertically symmetrical. The specifications of the robot are listed in Table II. TABLE II. SPECIFICATION OF ROBOT WITH MW-TRACK Length (mm) 350 Width (mm) 242 Height (mm) 110 Mass (g) 1050 Mass of a track belt (g) 200 Length of the track belt (mm) 240 Lap length of the track belt (mm) 508 Width of the track belt (mm) 22 sprocket outer diameter (mm) 84 Length of the shaft to tail end (mm) 160 Height of the grouser (mm) 10 IV", + " Tests on Simulated Environments In contrast, the robot with a 110 mm height continuous track could climb a 160 mm step (R of 145%) at the maximum. It was 2.9 times higher than that of the conventional wheeled robot. For all successful cases of the proposed robot, the grousers continuously scratched the corner of the step until one them hooked it. Afterward, the body was pulled with the support of the tail. In case of failure, the grousers were either not hooked at the corner or detached from the surface, following which the robot fell off. To confirm the motion of MW-Track on rough terrain, we tested it with the mobile robot shown in Fig. 2 on concrete blocks. The size of block was H 390 mm x W 190 mm x D 100 mm. The one of the blocks was lay flat and the other was piled on the first one with inclined. The procedure of the experiment was shown in Fig.7. The robot climbed the first block with same way as step-climbing (Fig.7, 1-2). Second, it got over the gap between the blocks without stacking by deforming the track belt and hook the grouser on the corner of the second block (Fig.7, 3-5). Then the robot got down the block from the front (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001741_eleco.2015.7394513-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001741_eleco.2015.7394513-Figure4-1.png", + "caption": "Fig. 4. Moto", + "texts": [ + " The model is nonline aerodynamic load of the propeller ( The parameterized model is g armature voltage and the state varia 6 CDE + C- The pulse width modulation (PW the armature voltage as given below % F GH IJK F L MN M %%3 2 (11) el approach is given below, nt of inertia of a rotor. 0 * (12) + 345 7 $%% & (13) ),.& (14) + ;O;<< (15) 0 ;O;== (16) ;== ? (17) med to be small, these body to the inertial frame. The ered small and they can be nonlinear equations of the (18) (19) (20) odel l be sufficient for most of the ar because of the nonlinear P 0 /). See fig. 4. r model iven below, where u is the ble x is the motor speed. + C! ! (21) M) duty can be converted to . QJRRS T GH IJK (22) 3. Quadrotor Platform Crazyflie quadrotor is used in this wor quadrotor that enables the implementation o to run on its onboard microcontroller communication link to a computer. It can b generic USB joystick. See fig. 5. To identify the motors, a special spee implemented to one of the motors. The mo PWM and the battery voltage are record exciting the motor giving throttle (in open-lo is converted to the armature voltage using voltage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003948_j.oceaneng.2021.108586-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003948_j.oceaneng.2021.108586-Figure2-1.png", + "caption": "Fig. 2. Relative position and orientation of guide ship A and approach ship B.", + "texts": [ + " The second phase is the key phase during which replenishment operations are conducted. Both ships must move with the same surge speed and heading angle. Considering the motions in surge, sway and yaw, the 3-DOF kinematic model of ship can be expressed as: \u03b7\u0307i = J(\u03c8i)\u03c5i with J(\u03c8i) = \u23a1 \u23a3 cos \u03c8i \u2212 sin \u03c8i 0 sin \u03c8i cos \u03c8i 0 0 0 1 \u23a4 \u23a6, i = A,B (1) where index i represents guide ship A and approach ship B. \u03b7i = [xi, yi,\u03c8 i] T and \u03c5i = [ui, vi, ri] T are the position vector and velocity vector of ship i, respectively. The coordinate system for ships A and B is shown in Fig. 2, where FXi, FYi and Ni denote the forces and moments acting on ship i. s1 and s2 are the transverse distance and longitudinal distance between them, respectively. The maneuvering modeling group (MMG) model is adopted as the dynamic model for both ships (Zhang et al., 2013): \u23a7 \u23aa \u23a8 \u23aa \u23a9 (mi + \u03bb11i )u\u0307i \u2212 (mi + \u03bb22i )viri = XHi + XPi + dXi (mi + \u03bb22i )v\u0307i + (mi + \u03bb11i )uiri = YHi + dYi (IZi + \u03bb66i )r\u0307i = NHi + NRi + dNi , i=A,B. (2) with \u23a7 \u23aa \u23a8 \u23aa \u23a9 XPi = (1 \u2212 tP)\u03c1wn2 Pi D4 Pi Kt NRi = \u2212 (1 + \u03b1H)xRi FNi cos \u03b4i FNi = 0", + " 5, the geometric relationship between the virtual trajectory and the trajectory of guide ship A can be expressed as: \u03b7V = \u03b7A + J(\u03c8A)L (14) where L = [lV cos(\u03b3V), lV sin(\u03b3V), 0] T. \u03b7V = [xV , yV ,\u03c8V ] T denotes the virtual trajectory, which should be tracked by the approach ship. The relative distance between the two ships is set according to different kinds of ships, cargoes and transporting equipment, which means that the values of lV and \u03b3V depend on the practical situation. It is obvious that s1 and s2 shown in Fig. 2 should be kept within a small range of lV sin(\u03b3V) and lV cos(\u03b3V) when performing the replenishment task, and the heading angles of the two ships should be kept consistent. The following errors are defined: \u23a7 \u23a8 \u23a9 es1 = s1 \u2212 lV sin(\u03b3V) es2 = s2 \u2212 lV cos(\u03b3V) e\u03c8 = \u03c8A \u2212 \u03c8B (15) The errors in Eq. (15) should be kept within a small range of zero. In Fig. 5, the desired heading angle of approach ship is obtained: \u03c8dB = 1 2 [1 \u2212 sgn(xV \u2212 xB)]sgn(yV \u2212 yB) + arctan yV \u2212 yB xV \u2212 xB (16) Based on the propeller speed nPB , rudder angle \u03b4B and measurable position vector \u03b7B = [xB, yB,\u03c8B] T of approach ship B, an ESO is proposed in this section to obtain the estimations of unmeasured speed vector \u03c5B = [uB, vB, rB] T and total uncertainties \u03b6 = [\u03b6u, \u03b6v, \u03b6r] T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000785_s10846-013-9973-9-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000785_s10846-013-9973-9-Figure4-1.png", + "caption": "Fig. 4 Testbed mechanism", + "texts": [ + " The parasitic moment vector applied from the testbed to the helicopter is denoted by Mtb. This is due to the damping effect of the cylinder, and its magnitude can be written as Mtb = \u2212rb \u0307 cos (\u03c0 2 \u2212 \u03b8 \u2212 \u03b1 ) = \u2212rb \u0307 sin(\u03b8 + \u03b1) (2) where, b is the damping coefficient of the cylinder. \u0307 is the rate at which the cylinder elongates and is calculated by taking time derivative of the loop closure equation in the complex plane for the planar mechanism \u201cOBC\u201d. The following vector relation holds for the mechanism shown in Fig. 4 l \u2212 r = h (3) The loop closure equation can be written in the complex plane as e j( \u03c02 \u2212\u03b1) \u2212 r e\u2212 j( \u03c02 \u2212\u03b8) = jh (4) where j = \u221a\u22121. Using Euler\u2019s formula results in (sin(\u03b1)+ j cos(\u03b1))\u2212 r(sin(\u03b8)\u2212 j cos(\u03b8)) = jh (5) Rearranging Eq. 5 results in sin(\u03b1)\u2212 r sin(\u03b8) = 0 (6) cos(\u03b1)+ r cos(\u03b8) = h (7) Since the testbed is designed to test the helicopter around hover, the angles \u03b8 and \u03b1 are small. Assuming small angles for \u03b8 and \u03b1 in Eqs. 6 and 7 results in \u03b1 \u2212 r\u03b8 = 0 (8) + r = h (9) Then, differentiating from Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000986_gt2014-26756-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000986_gt2014-26756-Figure2-1.png", + "caption": "FIGURE 2. CROSS SECTION OF THE TEST RIG [11]", + "texts": [ + " First, the experimental setup and the measurement techniques are briefly described. The method of identification of a regime change is then explained and the experimental results are shown. With this data, the impact of the different driving parameters is discussed and characteristic numbers are proposed that allow a prediction of the flow regime inside a bearing chamber. The experiments were carried out with the high speed bearing chamber test rig at the Institut fu\u0308r Thermische Stro\u0308mungsmaschinen at the Karlsruhe Institute of Technology. Figure 2 gives an overview over the test rig, which was also used in the same operational setup by Kurz et al. [11]. They also give a detailed description of the test rig, which will shortly be summarized in the following. The test rig consists of a shaft with an engine typical diameter, however much shorter, which is supported by two bearings. Next to each bearing there are bearing chambers, one of which was modified and tested for this investigation. The bearing next to the test chamber is lubricated with oil via under cage lubrication: Oil is injected through a nozzle at the inside of the hollow shaft (red line in Figure 2). Little radial holes in the shaft then guide the oil to the cage of the bearing. After having passed through the bearing, the oil flow splits and enters both of the adjacent bearing chambers. A scavenge pump is attached to each chamber and scavenges the air/oil mixture at a constant flow rate 2 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/80924/ on 05/01/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use from the bearing chamber" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001689_s11071-015-2356-y-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001689_s11071-015-2356-y-Figure4-1.png", + "caption": "Fig. 4 Demonstration of a stable point with respect to slipping, the trajectory returns to Case NR. Left panel effect of perturbation in time. Right panel projection of the trajectories onto the orthogonal space", + "texts": [ + " (59) By substituting (58) into (56), analytical formulae are obtained which determine a subset of stable points in XNR. For the stationary solution (46) of Case NR, \u03b21 = 0, \u03b22 = \u03c0 and we obtain \u03b3 (x\u2217, \u03b21) = \u2212 j+1 r j ( \u03b7R\u03c92 0 \u2212 g ) < 0, \u03b3 (x\u2217, \u03b22) = \u2212 j+1 r j ( \u03b7R\u03c92 0 + g ) < 0, (60) that is, (46) is stable with respect to slipping if \u03c90 > \u221a g \u03b7R . (61) In this case, the dynamics extinguishes the effect of the perturbation, and after awhile, the ball returns to the stationary solution of Case NR (see Fig. 4). However, if \u03c90 < \u221a g/(\u03b7R), then the stationary solution is unstable with respect to slipping and the dynamics repels the trajectories away from the stationary solution of Case NR (see Fig. 5). If only neutral stability is required then instead of (61), we get the same result as in (46), which was obtained from the requirement (13). By direct calculation from (27) and (56), one can check that this agreement is valid not only for the stationary solution but also for any point in XNR. The reasons behind this coincidence can be explained from the properties of the simple Coulomb friction model (see [12], p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003013_b978-0-08-102832-2.00026-8-Figure26.8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003013_b978-0-08-102832-2.00026-8-Figure26.8-1.png", + "caption": "Fig. 26.8 (A) Illustration of a polymeric disc, such as the one reported by Ismagilov et al. [67] that contains Pt on only one side of the structure. In this case the motion is triggered via bubble propulsion, that is, by the unidirectional expansion of O2 bubbles. The green arrows indicate the net propulsive force. (B) Schematic illustration of the motion triggered in a segmented Au/Pt wire via a self-electrophoresis mechanism. It involves an internal electron flow from the Pt segment to the Au section and a migration of protons in the surroundings of the wire.", + "texts": [ + " For example, they can be easily compartmentalized in different modules or segments (see Section 26.2), enabling a larger loading of multiple drug payloads [66]. Moreover, the versatility of their design is also an advantage to accomplish a specific task or application. Since the pioneering work of Ismagilov et al. in 2002 [67], where millimeter-scale polydimethylsiloxane (PDMS) discs with an integrated platinum rudder were observed to move autonomously in an aqueous hydrogen peroxide solution (Fig. 26.8A), much effort has been made to establish new catalytic-based engines at lower scales, that is, at the micro- and nanorealm. In the earlier investigation by Ismagilov et al., the structures propelled by the catalytic decomposition of hydrogen peroxide (H2O2) at the Pt rudder, which led to formation of oxygen (O2) bubbles (see Eq. 26.4) that exerted a propulsive force. While in this groundbreaking work, motion is mainly triggered by the unidirectional expansion of O2 bubbles (bubble propulsion); other asymmetric swimmers have also been reported, which exhibit different motion characteristics (vide infra)", + " In that study the asymmetric deposition of the catalyst (Pt) was realized based on previous investigations by Natan and coworkers [69], who described the fabrication of segmented nanowires via sequential electrochemical deposition (see Section 26.2). The authors described the motion of 2-\u03bcm-long platinum/gold (Au/Pt) wires with a diameter of 370 nm in 2%\u20133% hydrogen peroxide solutions. Surprisingly, and in sharp contrast to the bubble propulsion mechanism demonstrated by Ismagilov et al., the movement of Au/Pt wires was observed along the wires\u2019 axis but in the direction of the Pt segment (Fig. 26.8B). The authors suggested that the observed motion mechanism was the result of an interfacial tension gradient being continually reestablished on the surface of the moving object during the decomposition of H2O2. Likewise, Ozin and coworkers presented self-powered gold/nickel (Au/Ni) nanorods that were around 2-\u03bcm long and 200 nm in diameter [70]. In this case the authors stated that the catalytic decomposition of H2O2 occurred at the Ni segment and verified their claim by fabricating nanowires without a Ni component", + " From this extensive effort, it is arguably accepted that catalytic bimetallic micro- and nanowires propel in a hydrogen peroxide solution via self-electrophoresis [71]. This mechanism involves various concurrent chemical processes happening within a single segmented wire and its surrounding media. To explain this motion mechanism clearly, let us consider the Au/Pt wires reported by Mallouk et al. in 2004 [68]. An oxidation of the hydrogen peroxide occurred at the anode (Pt segment) yielding oxygen, protons, and electrons (see Fig. 26.8B and Eq. 26.4), with a reduction of oxygen and hydrogen peroxide at the cathode (Au segment) (see Fig. 26.8B and Eqs. 26.5, 26.6). In the Au section of the wire (Fig. 26.8B), protons and electrons are also consumed. For these two reactions to take place in the wire, an electron flux within the wire and a diffusion of protons from the anode to the cathode are necessary. The diffusion of protons can be explained by their electromigration through the double layer at the solution/wire surface interface that concomitantly generates an electric field via establishing a proton concentration gradient (Fig. 26.8B). Therefore it is believed that the electron flow and the corresponding proton diffusion facilitate the effective propulsion mechanism of bimetallic micro- and nanowires in the direction of the Pt segment (from the Pt section forward) when exposed to hydrogen peroxide [72, 73]. Note that this propulsion mechanism is apparently contradictory to that reported by Ismagilov et al. [67], where the unidirectional expansion of O2 bubbles on one side of the polymeric discs triggered the motion (Fig. 26.8A): Anode (Pt segment): H2O2 !O2 + 2H + + 2e (26.4) ode (Au segment): Cath O2 + 4H + + 4e ! 2H2O (26.5) ode (Au segment): Cath H2O2 + 2H + + 2e ! 2H2O (26.6) that other authors or some authors still consider that the motion in bimetallic Note nanorods and other catalytic nanoswimmers can also be established by changes in the liquid-vapor interfacial tension [68] or by changes in the viscosity close to the motor due to oxygen bubble generation [74]. The formation of concentration gradients around micro- and nanomotors can also lead to locomotion, a process that is also referred to as self-diffusiophoresis [75]", + " also demonstrated that the rotational motion of Pt/Si nanorods is triggered by the asymmetric deposition of the catalyst [65], which is the propelling force to push the nanorod forward from the deposited Pt layer. This result was further corroborated by designing L-shaped Pt/Si nanostructures where the direction of motion could be undoubtedly determined by the asymmetry of the shape of the swimmer. Indeed the authors showed that L-shaped Pt/Si structures moved predominantly in the direction opposite to the Pt end. Note that, in both studies, the Pt/Si and L-shaped Si/Pt nanostructures were composed of an insulator/metal interface where the electron flow indicated in Fig. 26.8B is hampered. These investigations corroborate the initial studies of theWhitesides\u2019 group where the catalytic decomposition of hydrogen peroxide on the catalyst surface is suggested as the main motive for generating the propelling force. Importantly, these results are also in accordance with the accepted interpretation that translational motion in bimetallic micro- and nanowires is achieved via a self-electrophoresis mechanism. What is missing in all of these first-rate examples of micro- and nanoswimmers is how to harness these systems with controlled motion capabilities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001308_iemdc.2013.6556236-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001308_iemdc.2013.6556236-Figure9-1.png", + "caption": "Fig. 9. Principal stress planes and flux density vector.", + "texts": [ + " Then, the eddy current and hysteresis losses are calculated by the post 1-D FEM from the time-variation in flux density at each 2-D finite element in order to consider the variation in the electromagnetic field along the thickness of electrical steel sheets. The effect of the stress on the core loss is taken into account by using the method reported in [3], in which the difference between the direction of the flux and that of the stress is considered in the core loss calculation. First, the flux density vector B calculated by (12) at each finite element and time step is decomposed into two components B1 and B2 along the principal axes of the stress, as shown in Fig. 9. Next, the 1-D nonlinear time-stepping FEM is carried out at each 2-D finite element in the core region by considering the principal axes, as follows: t A z A B z core VM 11 1 ,' (14) t A z A B z core VM 22 1 ,' (15) where A1 and A are the components of the magnetic vector potential along the principal axes, respectively; core is the resistivity of the electrical steel sheet. The analysis region is half the thickness of the electrical steel sheet. The total flux in this region is set to be identical to that of the element in the 2- D FEM by imposing following boundary conditions: )0,0(),( 0210D1 zz AAA (16) ) 2 , 2 (),( 12 2 21 2 D1 hBhB AA h z h z A (17) where h is the thickness of the sheet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001068_s1068375513030022-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001068_s1068375513030022-Figure6-1.png", + "caption": "Fig. 6. Dependence of surface roughness (Ra) on TE feed rate after deposition using different ESA methods: (1) with constant number of deposited layers (method 1) and (2) with constant energy amount (method 2).", + "texts": [ + " The change of the elemental composi tion of nano(micro)fibers should lead to the change of the surface properties. The elemental analysis of the total surface shows that the surface layer consists of the mixture of alumi num and tin oxides with different content of these ele ments, which depends on the modes of treatment. SURFACE ENGINEERING AND APPLIED ELECTROCHEMISTRY Vol. 49 No. 3 2013 ELECTROSPARK ALLOYING FOR DEPOSITION ON ALUMINUM SURFACE 185 One main peculiarity of the ESA is a very high rough ness of the surface. As is seen from Fig. 6, it depends strongly on the TE feed rate across the specimen. The highest roughness is observed at low TE feed rates. The interval values of Ra in Fig. 6 show the roughness change of a few specimens but not the roughness variation on the surface of the single one. Thus, e.g., if for the surfaces manufactured at V = 0.3 mm/s, a standard deviation for Ra at different points on a single specimen was from 3 to 6% of the measured value, for different specimens the deviation from the average values of Ra was up to 40% (Fig. 6). This again emphasizes (along with the data in Fig. 4) that the ESA is a process which is hard to control from the quantitative viewpoint. However, in spite of the spread in the experimental data, the treatment using different methods yields close results as to the roughness. In a wide range of the TE feed rates there are simi lar values of Ra for the both methods 1 and 2 (Fig. 6). Exceptions are in the experiments with the specimens 186 SURFACE ENGINEERING AND APPLIED ELECTROCHEMISTRY Vol. 49 No. 3 2013 AGAFII et al. obtained by method 2 in the \u201csparking\u201d mode (the experiments at V = 2 mm/s) (Figs. 6, 7). The results of the wear tests showed that the value of K (provided that the treatment is performed in mode I at a building up of coatings) can indeed exceed the unity (the rate of wear of the treated surface was 2\u2013 3 fold less that the rate of the counterbody wear, as given in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure8-1.png", + "caption": "Fig. 8. Temperature distribution with the spiral cooling duct at inlet flow rate 10 l/min and temperature 65oC and motor drive point (1100rpm, 101Nm).", + "texts": [ + " 869 Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 15,2021 at 16:48:44 UTC from IEEE Xplore. Restrictions apply. The resulting temperatures achieved when having a spiral cooling duct is quite like that of the wave cooling duct, as seen in the comparison in Table IV. Also the temperature distribution is similar, with the hottest region found in the end windings. The end region closest to the duct outlet is slightly hotter than on the other side. The difference depends on the inlet flow rate and temperature. This is seen in Fig. 8 for the spiral duct. See also more examples of resulting temperature distribution and of water and air velocity in Fig. 9 for the operating point (5000 rpm, 48 Nm). The air gap diameter is 160.4 mm (outer rotor) and 161.9 mm (inner stator). The speed 5000 rpm corresponds to a rotor peripheral speed of 5*80.2*\u03c0/60=21 m/s, so simulations agree with calculations, see Fig. 9c of air-gap velocity on the xy-plane (in the middle of the motor). The heat transfer coefficients of the wave and spiral cooling ducts are investigated for different flow rates, showing slightly higher values for the spiral duct, see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002621_icmic.2016.7804303-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002621_icmic.2016.7804303-Figure1-1.png", + "caption": "Fig. 1: Multi-rotor UAV in the fixed inertial frame", + "texts": [ + " The air drag on each rotor induces a reactive torque \u03c4 opposite to the direction of the propeller rotation and proportional to the square of the rotational speed: \u03c4i = d \u03c92 i (2) where d > 0 is the drag coefficient of the rotor propeller i and \u03c9i is the corresponding rotational speed. From Newton\u2019s second law, the equations of UAV motion can be written as: m \u03be\u0308/E = R Ft/B \u2212mg ez, R\u0307 = R sk(\u2126), I\u2126\u0307 = \u2212\u2126\u00d7 I\u2126 + \u03931 \u2212 \u03932 (3) with: 225 978-0-9567157-6-0 c\u00a9 IEEE 2016 m \u2208 R the total mass of the N -rotor UAV \u03be \u2208 R3 the position (x, y, z) of the center of mass of the system expressed in the inertial frame E = (ex, ey, ez), ez is the vertical axis directed upwards (see Fig. 1). R Ft/B the total thrust force expressed in frame E g the gravity constant \u2126 \u2208 R3 the angular velocity expressed in frame B sk(\u2126) the skew-symmetric matrix associated to \u2126 I \u2208 R3\u00d73 the inertia matrix with respect to frame B \u03931 \u2208 R3 the total active torques in frame B, \u03931 = (\u0393roll,\u0393pitch,\u0393yaw)T \u03932 \u2208 R3 the gyroscopic effects acting on the UAV R \u2208 SO(3) the rotation matrix from frame B to E, R = c\u03b8c\u03c8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03b8s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 \u2212s\u03b8 s\u03c6c\u03b8 c\u03c6c\u03b8 (4) with \u03c6, \u03b8, and \u03c8 represent Euler angles (roll, pitch, and yaw) relatively to frame E" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003141_s40430-020-02423-1-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003141_s40430-020-02423-1-Figure2-1.png", + "caption": "Fig. 2 Photograph of experiment set-up", + "texts": [ + " Contrast to USR, smoother surface may be obtained due to the roller acts on the shaft under the longitudinal ultrasonic vibration L along the radial direction and the torsional ultrasonic vibration T around the radial direction. In this work, the shaft made of 7050 aluminium alloy will be processed by TDUSR. Its chemical compositions (in wt%) are listed in Table\u00a01. Before TDUSR, the shaft is pre-treated by turning into a shaft with 45\u00a0mm in diameter and 180\u00a0mm in length. And its surface roughness Ra is Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:325 Page 3 of 9 325 1.96\u00a0\u00b5m, the microhardness is 500HL and the residual stress \u03c3 is 3.87\u00a0MPa. The experimental set-up is shown in Fig.\u00a02. The twodimensional ultrasonic surface rolling system is mounted on the lathe carriage of the CA6140B/A lathe. The roller with a diameter of 12\u00a0mm is made of YG8 cemented carbide. And the roller vibrates with a resonant frequency 19,800\u00a0Hz, longitudinal ultrasonic vibration amplitude L 9\u00a0\u00b5m and the torsional ultrasonic vibration T 4\u00a0\u00b5m. After processed, the surface microtopography was observed by the tool microscope MF-U (Japan). Meanwhile, the surface roughness Ra, the microhardness HL and the residual stress \u03c3 of the shaft were examined with the profilometer SURTRONIC 3 + (UK), Leeb hardness sclerometer MH-5 (P" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002938_1687814020908422-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002938_1687814020908422-Figure7-1.png", + "caption": "Figure 7. The meshing relationship of (a) external and (b) internal meshing types in planetary gear set.", + "texts": [ + " Furthermore, the instantaneous meshing power loss function of planetary gear set is obtained Pm(t)= Xn i= 1 mi sp(t)F i spn(t)v i sp(t)+mi pr(t)F i prn(t)v i pr(t) h i \u00f020\u00de where mi sp(t) and mi pr(t) refer to the friction coefficients of \u2018\u2018sun\u2013planet\u2019\u2019 gear pair and \u2018\u2018planet\u2013ring\u2019\u2019 gear pair at time t, respectively; Fi spn(t) and Fi rpn(t) are the mesh- ing forces of \u2018\u2018sun\u2013planet\u2019\u2019 gear pair and \u2018\u2018planet\u2013ring\u2019\u2019 gear pair at time t, respectively; vi sp(t) and vi rp(t) denote the relative sliding speeds of \u2018\u2018sun\u2013planet\u2019\u2019 gear pair and \u2018\u2018planet\u2013ring\u2019\u2019 gear pair, which are calculated as follows. The external and internal meshing types exist simultaneously in planetary gear set, and the meshing relationship of the two types is illustrated in Figure 7. In the figure, rbx denotes the radius of the base circle, ax refers to the rotating angle, K is the meshing point, rx denotes the equivalent radius at the meshing point K, M1M2 is the friction direction at meshing point K, and x=1, 2. Thus, the relative sliding speed between gear 1 and gear 2 can be obtained Because the planet gears are not only rotating around their own centers but they are also rotating around the sun gear center with the planet carrier, vsp and vpr can be obtained based on equation (12) vsp = (vs vc)rb2 tana2 (vp vc)rb1 tana1 \u00f022\u00de vpr = (vr vc)rb2 tana2 (vp vc)rb1 tana1 = (vc vp)rb1 tana1 vcrb2 tana2 \u00f023\u00de where vs, vc, vp, and vr represent the rotating speeds of sun gear, carrier, planet gear, and ring, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001401_12.2041268-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001401_12.2041268-Figure9-1.png", + "caption": "Fig. 9: The NIR image shows three distinct zones of the welding process: The capillary at the location of the laser welding spot is characteristic for deep penetration welding; the melt pool is determined by the molten metal which solidifies and forms the bead, which still emits thermal radiation according to its local temperature.", + "texts": [], + "surrounding_texts": [ + "The images are captures at a rate of 100 full frames per second, which is limited by the InGaAs camera. For image processing we use a personal computer with an optional frame grabber board. With this configuration every single image can be processed and evaluated individually and characteristic measurements can be determined. To emphasize specific weld features a \u201cpre-processing\u201d like generating a statistical image (chap. 3.2) may be helpful. The melt pool boundary usually can be determined from a step in grey value because the temperature of the liquid metal is approximately constant and the temperature drops after solidification 4 . Lower temperature results in a less thermal radiation. Of course this is quite simplified and is true only for an ideal grey body with constant emissivity 2 . But if an absolute temperature value is not required, the simple grey value evaluation delivers correct melt pool boundaries in many cases. (On the left hand side the image is attenuated by a 10%- neutral density filter to avoid saturation of the camera.)" + ] + }, + { + "image_filename": "designv11_34_0002344_s00170-016-9132-0-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002344_s00170-016-9132-0-Figure12-1.png", + "caption": "Fig. 12 Simulation model of the structural stage. a Boundary conditions for structural analysis. b The deformed shape of the joint specimen. c Testing force extracted from the simulation result", + "texts": [ + " In this simulation, 100 cooling stepswere used with a step increment of 5. Similar to the heating stage analysis, a convection heat transfer rate of 2 W/m2 and an ambient temperature of 20 \u00b0Cwere also used for the free surface (Fig. 11b). To obtain the testing force, a structural analysis of the infrared staking joint was performed. The testing tool and workpiece were modeled as perfectly rigid and plastic materials, respectively. The velocity of 100 mm/s in the z-direction was applied to the testing tool (Fig. 12a). A testing stroke of 6 mm was used to reach the final step. A constant shear friction of 0.42 was assumed to obtain the best results in view of the testing forces. During the structural simulation, the load produced by the relative displacement between the testing tool and the workpiece is increased. A further increase in the load increases the internal stresses which reach the material ultimate tensile. After each structural analysis, the testing force was observed by extracting the simulation result (Fig. 12c). Figures 13 and 14 show the respective temperature history and deviation for these three heating times. The final workpiece temperatures for 8, 14, and 20 s were approximately 49, 59.9, and 66.7 \u00b0C, respectively (Fig. 13a\u2013c), and the temperature deviations were 3.2, 2.4, and 1.8 \u00b0C (Fig. 14a\u2013c). The small temperature deviations between the two points (P1 and P2) indicated that the final temperature distribution was uniform. It can be observed from the figures that the temperature deviation gradually decreased with increasing heating time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003552_j.sna.2020.112347-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003552_j.sna.2020.112347-Figure2-1.png", + "caption": "Fig. 2. Experimental setup: (a) hollow sphere filled with a ferrofluid emulsion freely suspended in water; (b) hollow sphere filled with a ferrofluid emulsion on the solid rface o", + "texts": [ + " Thus, the internal video a i m e a r I i l t 3 r fi i m f n f d u g t p e t r t nonmagnetic surface surrounded by water; (c) nonmagnetic sphere floats on the su reservoir filled with water; 3 \u2013 nonmagnetic plastic sphere; 4 \u2013 Helmholtz coils; 5 \u2013 rotations of microdrops can produce a torque acting on a whole emulsion sample due to the viscous coupling. To measure the torque, the emulsion sample was placed into a hollow plastic sphere which is neutrally buoyanced in the surrounding water. The sphere external diameter was 9 mm and internal diameter was 8.1 mm. The water density was adjusted by dissolving sodium chloride in it to obtain the desirable value. The sphere with the emulsion inside was positioned sufficiently far away from any boundaries so their influence can be neglected. The sketch of the experimental setup used is shown in Fig. 2(a). Under the action of a uniform rotating magnetic field the sphere rotation was observed and recorded through the camera. In this case the sphere corotates with the magnetic field. The sphere stationary rotation frequency was measured, from which the toque acting on the sphere can be easily determined. The macroscopic torque generated inside the sphere under a rotating magnetic field due to the internal rotations of microdrops is balanced by a viscous torque acting outside from the surrounding fluid: M = 8 1R3\u02dd, where 1 is the dynamic viscosity of a fluid surrounding the sphere, R is the sphere radius, is the sphere angular velocity", + " The magnitude of the resulting torque in ferrofluid emulsions is higher than that in pure ferrofluids of equal or even higher magnetic content. The observed macroscopic torque can be used for the controlled transport purposes. To examine the transport possibilities of ferrofluid emulsion the sphere with emulsion inside was placed on the flat bottom of the rectangular glass reservoir filled with water. The water reduced but did not fully compensated the gravity force exerted on the sphere in this case (Fig. 2(b)). It was observed that the sphere is propelled forward under the action of magnetic field rotating in a vertical plane. f t e f ferrofluid emulsion. 1 \u2013 hollow sphere filled with a ferrofluid emulsion; 2 \u2013 glass camera. Another manifestation of internal rotations of microdrops in ferrofluid emulsion is the actuation of macroscopic bodies mmersed in the emulsion. Particularly, we have considered a non- agnetic plastic sphere which is floating on top of a ferrofluid mulsion layer approximately half submerged (Fig. 2(c)). Under the ction of a magnetic field rotating in a horizontal plane the sphere otation in the direction opposite to the field rotation was observed. t has been found that under the action of a magnetic field rotating n a vertical plane the sphere moves along the surface of emulsion ayer. This time the direction of the sphere motion is opposite to he direction the field is rolling. . Experimental results The macroscopic torque generated in a finite volume of ferofluid emulsion was measured as a function of the magnetic eld strength and frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000216_gt2010-23035-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000216_gt2010-23035-Figure2-1.png", + "caption": "Fig. 2 Deformation in balls and races", + "texts": [ + " The foregoing procedure using analogous plane treatment simplifies the procedure when it is needed to add further details of the system such as internal damping and bearing stiffness, damping and support flexibility. Assuming perfect rolling of balls on the races, the varying compliance frequency of a ball bearing is given by \u03c9vc = \u03c9rotor * BN where BN = { (Ri /(Ri + Ro)}* Nb (6) The BN number depends on the specifications and dimensions of the bearing. When centre of the inner race moves from O to O' as indicated in Fig. 2, and radial displacement in any direction \u03b8, due to displacement (O O\u2019) is more than the radial internal clearance (\u03b30); there is interference between inner race and balls. Consequently, the elastic deformation of balls and races takes place. In Fig.2, (x, y) denote the displaced position (O') of the centre of the inner race. From simple geometrical analysis, the expression for radial elastic contact-deformation (\u03b4(\u03b8i)) of ball-races at angular position \u03b8i is 0( ) ( )i i ixcos ysin\u03b4 \u03b8 \u03b8 \u03b8 \u03b3= + \u2212 (7) In case of \u03b4(\u03b8i) > 0, ball at angular position \u03b8i is loaded giving rise to restoring force with nonlinear characteristics because of Hertzian contact deformation [19]. If \u03b4(\u03b8i) < 0, the ball is not in the load zone (\u03b8LZ), and restoring force from the ball is set to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003466_s106836662004008x-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003466_s106836662004008x-Figure2-1.png", + "caption": "Fig. 2. The scheme of the friction chamber: (1) input (purified air); (2) the friction unit; (3) the disk; (4) the friction chamber; (5) output (air with wear particles).", + "texts": [ + " 347 Friction-wear characteristics were studied using the T-11 friction machine of the friction lining (finger) type: a disk that allows measuring, recording, and visualizing experimental results, namely friction force, linear displacement (wear), and temperature of the friction lining. A general view of the T-11 friction machine complete with equipment is shown in Fig. 1. The controlled parameters of the T-11 friction machine are the rotational speed of the disk and the load on the friction lining. For ecological evaluation of friction pairs, the friction unit (friction lining, the disk) was installed into the friction chamber, a diagram of which is shown in Fig. 2. From the compressor, air passes through a filter system and, in purified form, enters the friction chamber. During friction, wear particles are released, which are picked up by the air f low in the friction chamber. The air leaving the friction chamber with wear particles is analyzed using the TSI NanoScan SMPS Nanoparticle Sizer 3910 (nSMPS) and TSI Optical Particle Sizer 3330 (OPS). nSMPS classifies particles sized 0.01\u20130.42 \u03bcm into 13 groups with a frequency of once per minute. OPS classifies particles sized 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000762_gt2013-95442-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000762_gt2013-95442-Figure5-1.png", + "caption": "FIGURE 5. Tested seal arrangement (bst 13)", + "texts": [ + " The maximum pressure is limited by the maximum mass flow of the test rig or the pressure difference of the two independent exhaust sides, resulting in a axial thrust which is limited by the axial bearing. TABLE 2. Parameters of the speed variations Inlet pressure Speed Steps 25 bar 2.000 \u2014 10.000 rev/min 2.000 rev/min TABLE 3. Parameters of the pressure variations Test no. Inlet pressure Speed Steps 1 0 \u2014 maximum bar 10.000 rev/min 5 bar 2 0 \u2014 maximum bar 2.000 rev/min 5 bar Tested Seals & Arrangements The tested seal arrangement (bst 13) was a four seal multi stage arrangement with different types of core elements, consisting of inclined and straight bristle packs as shown in FIGURE 5. As it can be seen in FIGURE 5 and 6, the first and the second seal are combined to a tandem seal design, the third and fourth seal are used as single seals. As shown in FIGURE 5, only the first seal has an axial inclined bristle pack. TABLE 4 lists the parameters of all tested seals. The bore and bristle diameter of the first seal is different to the others, in order to realise a new tandem design with a higher stiffness compared to the steam tests before. In order to analyse an assumed performance influence of contoured back plates, a grooved back plate design [8] was tested at seal no. 3. The grooves are contrary to the bristle lay angle but also arranged in 45\u25e6 with a depth of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002949_s12541-019-00277-9-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002949_s12541-019-00277-9-Figure1-1.png", + "caption": "Fig. 1 New type of conical involute gear", + "texts": [ + " Although considerable research has focused on the DTE and a time-varying mesh stiffness, little focus has been given regarding the effects of the reciprocating axial movement of the driven gear to eliminate time-varying backlash combined with the two-sided mesh stiffness excitation on the DTE of NTCIG. Moreover, as the NTCIG needs to be reversed frequently in locating and tracking precision transmission systems, this paper places significant emphasis on the twosided tooth mesh analysis through the DTE of the gear pair based on the angular position detection of the anti-backlash strategy under varying load and speed excitations. The NTCIG tooth thickness varies due to differences in the tangential modifications and similarities in the radial modifications, as shown in Fig.\u00a01. Compared with the conventional variable tooth thickness gear, each transverse plane of the NTCIG has identical dedendum and addendum circles. Therefore, gear interference can be avoided during the axial movement of the driven gear to eliminate backlash. The coordinate system is built based on the rotation center of the driving gear. In Fig.\u00a02, the rotational angle \u03b8 of the driving gear is defined as the angle between the symmetrical line of the driving gear tooth and the horizontal plane. The NTCIG is divided into N equal pieces along the direction of the gear width b" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003915_j.triboint.2021.106876-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003915_j.triboint.2021.106876-Figure6-1.png", + "caption": "Fig. 6. Schematic view of the pinion assembly; tapered roller bearings (REB), pinion gear and seal.", + "texts": [ + " Due to the differences in size and oil volume, and to the power loss in the tandem tapered roller bearings, it is not possible to keep the temperature constant in both test chambers. So, it was decided to control and keep constant the temperatures in the Gear chamber, when the operating conditions allow that. Later on, those oil sump temperatures in the Gear chamber will be used for the tests with the complete axle box and with the crown gear alone. The tests were performed under dip lubrication, and the oil level was set between alub = 20 mm up to 24 mm, as shown in Fig. 4. This way it was possible to measure the pinion assembly torque loss (TP.asb V0 , Fig. 6), that includes the pinion gear churning (TP.g VZ0), tapered roller bearings (TP.rb VL0) and seal torque loss (TP.sl VD ), as expressed in equation (1). TP.asb V0 = TP.g VZ0 + TP.rb VL0 + TP.sl VD (1) The starting torque was measured using the torque wrench shown in Fig. 5. In order to properly measure the starting torque promoted by the pinion assembly tapered roller bearings, the crown assembly and the oil sump were removed (the bearings have been previously oiled) and the J.A.O. Cruz et al. Tribology International 157 (2021) 106876 differential was at rest (still)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003124_j.matpr.2020.04.239-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003124_j.matpr.2020.04.239-Figure7-1.png", + "caption": "Fig. 7. Mode shapes of spindle supported by radial PMBs.", + "texts": [], + "surrounding_texts": [ + "Ansys rotor dynamic and harmonic response tool provides a complete set of environment to study the unbalance force response and dynamic characteristics of spindle bearing systems. In the Ansys harmonic response tool, rotor dynamic results are incorporated to characterize the rotating structures. Three dimensional FEA model of the drilling spindle bearing system is generated and essential motion and load constraints are applied using rotor dynamic Ansys workbench tool. Tangential and axial movements are restricted at element 1 to assist simply support boundary condition by ignoring the shear effect; also rotation about y axis is achieved through radial motion of the spindle. The entire spindle bearing system is meshed using higher hexa mesh as shown in Fig. 5. Rotor dynamic and harmonic response analysis of spindle is carried for determining the natural frequencies, mode shapes, critical speeds, harmonic frequencies and peak amplitudes. Stability analysis is carried out by plotting Campbell diagram," + ] + }, + { + "image_filename": "designv11_34_0003274_i2015-15036-1-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003274_i2015-15036-1-Figure1-1.png", + "caption": "Fig. 1. a) Schematic diagram of the experimental micropipette deflection setup. b) Optical microscopy image of the bending of a young adult worm. The insets show a zoom in of the pipette deflection (\u0394x = 11.4 \u00b1 0.1 \u03bcm) between the points of wormsupport contact (bottom) and the end of the bending part of the experiment (top).", + "texts": [ + " Here we present a detailed experimental study of the viscous relaxation of the nematode and apply a wellknown viscoelastic model to describe the worm material. We will show that a pure Newtonian fluid can not correctly capture the viscous component responsible for relaxation, but that the implementation of a complex, power-law fluid is necessary to understand its relaxation. We find that the worm is strongly shear thinning and quantify the viscoelastic properties of young adult and adult C. elegans nematodes. Micropipette deflection was used to measure the material properties of C. elegans as shown in the schematic illustration of fig. 1(a) and in the optical microscopy image of fig. 1(b). In this technique, the deflection of a long (1\u20132 cm) and thin (\u223c 20\u03bcm) glass micropipette, which acts as force measuring spring, is force-calibrated and used as a sensor capable of measuring forces down to the nN range. This technique was previously introduced to probe the elastic properties of C. elegans by simply bending the worm [20], and has since then been used to study the active dynamics of the nematode crawling on a gel surface [10] as well as swimming in an infinite fluid [6] and close to solid boundaries [7]", + " An anesthetized worm was held by the pipette at the point of the vulva by applying suction, and centered between two simple supports. As in our previous experiments [20], the support was made of a thicker (\u223c 50\u03bcm) micropipette curved into the shape of a U and mounted on the opposite side of the chamber to a motorized translation stage. The worm could then be bent by pushing it into the gap between the two sides of the U-shaped support, which was done by moving the support (xu, from left to right in fig. 1) with a constant speed (vu) towards the worm. This causes a deflection (x) of the pipette to the right, a springlike force (F = kpx) to the left and the worm to bend (y = xu\u2212x = vut\u2212x). The experiments were run using an in-house LabView code controlling the motor and camera. All wild type (N2) C. elegans nematodes used in this work were young adults or adults. The worms were acquired from the Caenorhabditis Genetics Center and were cultivated according to standard methods [28] on Escherichia coli (OP50) nematode growth media (NGM) plates at 20 \u25e6C" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003684_012021-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003684_012021-Figure5-1.png", + "caption": "Figure 5. Detection results of second stage network", + "texts": [ + " In strategy 5, K-means clustering algorithm is used to optimize the initial anchor frame, AP value is increased by 1.3%, and detection rate is slightly improved. The above experiments show that the improved Faster RCNN strategy proposed in this paper can effectively improve the comprehensive detection performance of the target in this paper. 4.3.2 Pin defect identification. Import the CKPT model file generated by the training, input the test set for testing, and part of the test results are shown in Figure 5(the mark in the figure about pin_defect_0x01 means falling off, pin_defect_ 0x02 means prolapses), and diagnose whether the pin in the fastener falls off or prolapses. UPIOT 2020 Journal of Physics: Conference Series 1659 (2020) 012021 IOP Publishing doi:10.1088/1742-6596/1659/1/012021 Because the two-stage network will be called repeatedly, it requires higher speed. The increase of training and detection scale will reduce the detection rate, while the small training and detection scale may lead to the network under fitting and reduce the detection accuracy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000658_2015-01-0610-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000658_2015-01-0610-Figure5-1.png", + "caption": "Figure 5. Two-stage gear drive", + "texts": [ + " Although the stiffness matrix which contains variable meshing stiffness of two-stage gear pairs can be get mentioned above, it is not convenient for the analysis because the stiffness matrix is also variable. So modeling restoring torques is applied to simulate variable meshing stiffness of two-stage gear pairs. Transmission error is defined as the difference between ideal and real displacements of driven gear in the gear pair. Firstly, the transmission errors of 1st gear and final drive gear pair are given in Equation (21) and (22) respectively, and the two-stage gear drive is shown in Figure 5. (21) (22) Then, the restoring torque vector of variable meshing stiffness is introduced in order to facilitate the modeling of gear variable meshing stiffness. Ignoring the variable meshing stiffness in stiffness matrix of the system, the restoring torque vector which includes the restoring torques on the both pinion and wheel are added to the external torque vector of the system (on the right of Equation (12)). The governing equation of branched torsional vibration model which includes two-stage gear variable meshing stiffness is yielded in Equation (23)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001446_ijista.2014.059300-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001446_ijista.2014.059300-Figure2-1.png", + "caption": "Figure 2 The TREX600 helicopter used in the system identification experiment with instrumentation equipment fitted between fuselage and landing gear (see online version for colours)", + "texts": [ + " The helicopter model was selected due to its sufficient payload capacity, great manoeuvrability and low cost replacement parts. It was equipped with a standard Bell-Hiller stabiliser bar on the main rotor, which improves handling characteristics for human pilots by increasing the damping on the pitch and roll responses. Furthermore, TREX600 was also equipped with a high efficiency high torque brushless motor that allows the helicopter to carry about 2 kg payloads with an operation time of about 15 m. The basic UAV platform shown in Figure 2 had been modified to make room for installing necessary electronic equipments which gathers flight data for dynamic modelling and control system design. Some key physical parameters of TREX600 RC helicopter are given in Table 1. The flight test was conducted on the UAV helicopter platform in calm weather conditions. Different flight manoeuvres were conducted to excite the desired dynamic of interest. For example, after the helicopter reached steady and level condition, the yaw dynamics was excited using only tail collective pitch command while other input commands were used to balance the helicopter in such a way to make the vehicle oscillate roughly around the operating point of interest" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003911_s42417-020-00269-4-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003911_s42417-020-00269-4-Figure6-1.png", + "caption": "Fig. 6 Common clutch disc-damping elements", + "texts": [ + "3 Z\u0308 \u23ab\u23aa\u23ac\u23aa\u23ad + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 C12 \u2212C12 0 0 \u2212C12 C12 + C23 \u2212C23 0 0 \u2212C23 C23 0 1 \u2212 z \ud835\udeff sign \ufffd ?\u0307?2 \u2212 ?\u0307?1 \ufffd \u2212 \ufffd 1 \u2212 z \ud835\udeff sign \ufffd ?\u0307?2 \u2212 ?\u0307?1 \ufffd\ufffd 0 1 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a7\u23aa\u23a8\u23aa\u23a9 ?\u0307?1 ?\u0307?2 ?\u0307?3 Z\u0307 \u23ab\u23aa\u23ac\u23aa\u23ad + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 k12 \u2212k12 0 \u2212 Fmax \ud835\udeff \u2212k12 k12 + k23 \u2212k23 Fmax \ud835\udeff 0 \u2212k23 k23 0 0 0 0 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a7 \u23aa\u23a8\u23aa\u23a9 \ud835\udf031 \ud835\udf032 \ud835\udf033 Z \u23ab \u23aa\u23ac\u23aa\u23ad = \u23a7\u23aa\u23a8\u23aa\u23a9 \ud835\udf0f 0 0 0 \u23ab \u23aa\u23ac\u23aa\u23ad 1 3 In general, a clutch disc has a set of helical compression springs. A single pair of springs is shown in Fig.\u00a05, with cylindrical springs of two different diameters. Four pairs of springs are shown in Fig.\u00a06 inside the slots of the damper, as commonly found in clutch discs. The equivalent stiffness for these springs mounted in parallel results in stiffness k12. Tuning tests are carried out in the automotive industry to evaluate the vehicle by measuring the angular acceleration of the drivetrain components under five conditions (idle, creeping, drive, coast and engine start/stop). In this study, the drive condition was analyzed, which is the most common and severe condition of the powertrain. The engine torque flows along the entire driveline to the wheels during vehicle acceleration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003379_acc45564.2020.9147606-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003379_acc45564.2020.9147606-Figure2-1.png", + "caption": "Fig. 2: Diagram of the path and goal point. The Xp \u2212 Yp coordinate system is such that the Xp axis is tangent to the path point closest to the goal point. The red line represents a sample path and \u03b4 is the lateral offset to account for serpentine motion of the sleigh.", + "texts": [ + " This leads to a system of 11 nonlinear equations with 11 unknowns (u0, Au1, Bu1, Au2, Bu2, Aw1, Bw1, Aw2, Bw2, \u03c40, A). Since the equations are nonlinear we use Newton\u2019s method to solve them numerically to calculate the input required to produce the desired motion. In the absence of \u03c40 the average motion is along a straight line and \u03c90 will be zero. The reference angle \u03b8r is of course known in the preceding calculations. This reference angle will be derived from the path tracking algorithm. The schematic fig. 2 shows the geometric setup of vector pursuit in a plane. 5258 Authorized licensed use limited to: University of Wollongong. Downloaded on August 10,2020 at 06:44:26 UTC from IEEE Xplore. Restrictions apply. To calculate the reference angle required for the sleigh to track the path we employ a vector pursuit path tracking problem. First we define the coordinates and orientation of the path at the goal point to be (xt, yt, \u03b8t). The goal point is defined to be one look ahead distance L from the sleigh at any given time. The desired reference angle is then \u03b8r = sign(yt\u2212 (y\u2212 b sin(\u03b8))) cos\u22121 (xt \u2212 (x\u2212 b cos(\u03b8)) L ) which is the angle of the line connecting point P of the sleigh to the goal point as shown in fig. 2. Note that if the goal point is always on the path, the serpentine motion of the sleigh will cause oscillations in the reference angle. Such oscillations are undesirable as the harmonic balance assumes the reference angle to be either constant or linearly increasing and the harmonic balance solutions would not be valid when the reference angle is oscillating. In order to take this into account we require that the goal point executes serpentine motion around the path similar to the sleigh. This ensures that in the ideal case when the sleigh is following the path, \u03b8r remains constant or linearly increasing as necessary", + " Serpentine motion of the target can be prescribed by assuming ideal limit cycle motion for the target. We define \u03b4(t) to be the signed distance between the sleigh and its average path for t \u2208 [0, T ] where T = 2\u03c0/\u2126 is the time period. The goal point is then calculated using the following procedure 1) Find the point on the path closest to the sleigh 2) Beginning from this point search along the path for the first point that is a distance L from the sleigh 3) Find the point \u03b4(mod(t, 2\u03c0)) away from the above path point along the Yp direction as shown in Figure 2 and set this to be the goal point The effectiveness of the path tracking algorithm was tested with numerical simulations. In this section we show the sleigh\u2019s ability to track both linear and circular paths using the proposed algorithm. In each case we specify a velocity of v\u0304d = 0.2 and for the circle we chose a radius of R = 10. We define the normalized error in the velocity of the sleigh to be ev = v\u0304 \u2212 v\u0304d v\u0304d Figure 3 demonstrates the convergence of the proposed control law for a set of parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure1-1.png", + "caption": "Figure 1. The relationship of pitch curve of curve face gear pair: (1) the pitch curve of non-circular gear, (2) the pitch curve of curve-face gear and (3) the planar pitch curve of rack.", + "texts": [ + "edu.cn at The University of Melbourne Libraries on June 5, 2016pic.sagepub.comDownloaded from tooth contact analysis method with experiment. These results provide the theoretical support for further theoretical analysis and practical application of curve gear pair. Tooth surfaces of curve-face gear pair The generation of the pitch curve of curve-face gear can be established as follows: expanding the curve-face gear from cylindrical surface into a rack and it can result in a planar curve as shown Figure 1. As shown in Figure 1, point P is the meshing point of curve-face gear pair (not shown in Figure 1). Point P0 is on the pitch curve 1 with the corresponding meshing point P00 on the pitch curve 2, expanding point P000 is point P00 shown on planar curve 3 with rotating angle 1. According to the spatial meshing theorem, the velocity at tangency point P of curveface gear and non-circular gear must be equal to each other in the process of meshing, that is r\u00f0 1\u00de !1 \u00bc R !2 r\u00f0 1\u00de \u00bc a\u00f01 e2\u00de 1 e cos\u00f0n1 1\u00de 8< : \u00f01\u00de According to the theorem of meshing tooth surface, the transmission ratio of curve-face gear pair can be shown as follows i12 \u00bc " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002135_amc.2016.7496392-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002135_amc.2016.7496392-Figure11-1.png", + "caption": "Fig. 11. Force direction", + "texts": [], + "surrounding_texts": [ + "To simplify the equilibrium problem, it assume that the center angles of hold cones are zero degree in this paper. It means each force can be generated to only one direction on the corresponding hold . Suppose the case n = 3. If the hold positions, mass, and the force directions are assumed as x1 = \u22120.05;y1 = \u22120.15;z1 = 0.38; x2 = 0.01;y2 = 0.20;z2 = 0.40; x3 = 0.02;y3 = 0.25;z3 = \u22120.20; m = 3.0kg; g = 9.8m/s2, \u03d51 = \u03c0/8;\u03b81 = \u03c0/2 \u2217 3/5; \u03d52 = \u2212\u03c0/10;\u03b82 = \u03c0/2 \u2217 6/5; \u03d53 = \u2212\u03c0/6;\u03b83 = \u03c0/2 \u2217 7/5, The magnitude of forces F = [F1, F2, F3] T = [5.67, 41.2,\u221217.7]T N can be calculated, COM position P = [x, y, z] T = [0.182, 0.0156, 0.00]T m is also calculated but the amplitude of z component are arbitrary. Moreover, suppose \u03d51 = \u03c0/6;\u03b81 = \u03c0/2 \u2217 3.5/5\u3001\u03d52 = \u2212\u03c0/10;\u03b82 = \u03c0/2 \u2217 6.5/5; \u03d53 = \u2212\u03c0/6;\u03b83 = \u03c0/2 \u2217 7/5, The magnitude of the force is changed to F = [F1, F2, F3] T = [12.4, 51.7,\u221234.3]T N, and the position of COM is also moved to P = [x, y, z] T = [0.337,\u22120.182, 0.00]T m. It means equilibrium point can be changed with force direction. Note that not only COM equilibrium point but also magnitude of force are changed. Then robot controller needs decision of whether force direction is inside of hold cone or not." + ] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure18-1.png", + "caption": "Figure 18. Experimental setup to measure the dynamic frictional resistance of each gear type, including that with flat passive rollers", + "texts": [ + " The X-axis of the graph of Fig. 17 is the pressing force from the force gauge on the right. The average values after the experiments had been repeated 10 times were plotted on the graph, and the error bars show the standard deviation. The gear with flat passive rollers showed good smooth sliding performance. The static frictional resistance is always the minimum among the three gear types, and the t-test also showed significant difference in each case. In this experiment, the omnidirectional gear in Fig. 18 was driven by the same geared AC motor used in Chapter V. Its torque was kept constant (0.006 Nm) and the gears slid on the omnidirectional gear by the same distance of 56 mm. The result of the experiment is shown in Fig. 19. The purple horizontal line of \u201cFree slide\u201d is the average value of the omnidirectional gear\u2019s sliding repeated 10 times without any pressing force from the force gauge on the right. Even though the dynamic frictional resistance increased in the experiments of the spur and crowning gears, the dynamic frictional resistance of the gear with flat passive roller was almost constant and as small as the value of \u201cFree slide\u201d, hence confirming the advantage of the gear with flat passive rollers as compared to dynamic frictional resistance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002875_s42496-020-00033-7-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002875_s42496-020-00033-7-Figure12-1.png", + "caption": "Fig. 12 Trend of the torque correction term (feedforward term)", + "texts": [ + " The combination of the HOILC and FB controller has proved to be successful in enhancing the precision of the attitude tracking when performing repetitive manoeuvres. 1 3 As mentioned in Sect.\u00a03.1, the control input is composed of two terms: a feedback and a feedforward signals deriving from Eq.\u00a0(9). The two components have been plotted in Figs.\u00a012 and 13 both with respect to time and iteration (orbit). As the feedback term decreases iteration by iteration (see Fig.\u00a013), the feedforward term gradually increases (see Fig.\u00a012). However, the corrective term uC in Eq.\u00a0(9) (proportional to the angular position and velocity errors) will be zero when the desired trajectory is reached. Such a condition guarantees the feedforward torque does not increase indefinitely, but it remains constant for further iterations (as can be noticed in the last iterations reported in Fig.\u00a012). This means that the controller has learnt the amount of supplementary torque to add to the standard PD controller to track the trajectory in presence of disturbances given by elastic panels and sloshing masses. The left solar panel tip displacement is shown in Fig.\u00a014. The order of magnitude of the maximum deflection is 20\u00a0cm, but it is expected if considering the assumptions made in Sect.\u00a02.2. The reported tip displacement would correspond to a roughly half-degree relative rotation between two adjacent sections (if the array segments were considered to be rigid)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure16-1.png", + "caption": "Fig. 16 Drive system integrated inside the RL2 Hand\u2019s structure", + "texts": [ + " Cab\u00e1s Ormaechea These three sensors are located in the states mechanism selector, the MultiState Actuator, which interacts with the palm and is taken out to cover the ASIBOT contacts, when the Main Platform is removed. The extraction system of the RL2 Hand fingers keeps almost intact its functioning philosophy because the results obtained at the RL1 Hand has been successful. In this particular case, the design has been only focused on the energetic performance of the energy transformer. Pulleys have been located in all the tendons way avoiding them to pass over surfaces that can create unnecessary friction. In Fig. 16, it is possible to see the changes done for this last design and a detail of the pieces that compound it. The final result of the design is shown in Fig. 17. Robotic Hands 37 Ambrose R, Aldridge H, Askew R, Burridge R, Bluethmann W, Diftler M, Lovchik C, Magruder D, Rehnmark F (2000) Robonaut: NASA\u2019s space humanoid. IEEE Intell Syst Appl 15 Balaguer C, Gim\u00e9nez A, Jard\u00f3n A, Correal R, Cabas R, Staroverov P (2003) Light weight autonomous robot for elderly and disabled persons\u2019 service. Int Conf Field Serv Rob Balaguer C, Gim\u00e9nez A, Jard\u00f3n A, Cabas R, Correal R (2005) Live experimentation of the service robot applications elderly people care in home environments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.1-1.png", + "caption": "Figure 9.1 A Puma manipulator in its side and top views.", + "texts": [ + " with similar types and arrangements of the joints. This chapter also includes some conceptual manipulators that do not have generic names. They are not necessarily encountered in practice. However, they have certain kinematic features that make them interesting and instructive to study. Moreover, this chapter contains examples of deficient and redundant manipulators as well in order to demonstrate the particular complications that occur in their inverse kinematic solutions. A Puma manipulator is shown in Figure 9.1. Its generic name comes from the acronym PUMA that stands for \u201cProgrammable Universal Manipulation Arm.\u201d It was developed originally by Victor Scheinman at the Unimation Robot Company. Typical versions of Puma manipulators are the Puma-560 and Puma-600. A Puma manipulator comprises six revolute joints. So, it is designated symbolically as 6R or R6. Its wrist is spherical. Its joint axes are also shown in Figure 9.1 together with the relevant unit vectors. Its kinematic details (the joint variables and the constant geometric parameters) are shown in the line diagrams in Figure 9.2. The line diagrams comprise the side view, the top view, and two auxiliary views that show the joint variables that are not seen in the side and top views. The significant points of the manipulator are as follows: Kinematics of General Spatial Mechanical Systems, First Edition. M. Kemal Ozgoren. \u00a9 2020 John Wiley & Sons Ltd. Published 2020 by John Wiley & Sons Ltd" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001192_j.apsusc.2014.09.063-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001192_j.apsusc.2014.09.063-Figure1-1.png", + "caption": "Fig. 1. Scheme of the biosensor: 1 \u2013 rubber ring, 2 \u2013 layer of enzyme (bioreceptor) immobilized onto PVA-T, 3 \u2013 electrode layer (transducer) consisting of carbonaceous m a e", + "texts": [ + " Working electrode (biosensor) was designed by echanically attaching and fixing the membrane containing immoilized enzyme to the surface of the electrode. The enzyme was mmobilized on individual flexible support of 0.1% polivinylalcool coated terylene (PVA-T). Adsorption of 2 l of PQQ-GDH to the upport (\u00d8 2 mm) was the method for the immobilization of the nzyme. The prototype amperometric biosensor was constructed y coupling enzymatic membrane consisting of PQQ-GDH and PVA- with electrode surface consisting different oxidized graphites or ristine graphite. The sensor construction is presented in Fig. 1. .3.2. Electrochemical measurements Chronoamperometry measurements were performed using Please cite this article in press as: J. Razumiene, et al., Nano-structur Surf. Sci. (2014), http://dx.doi.org/10.1016/j.apsusc.2014.09.063 n electrochemical system (PARSTAT 2273, Princeton Applied esearch, USA) with a conventional three-electrode system comrised of a platinum plate electrode as auxiliary, a saturated g/AgCl electrode as reference and working electrodes (\u00d8 2 mm) PRESS Science xxx (2014) xxx\u2013xxx 3 based on carbonaceous materials" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003190_s12204-020-2195-y-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003190_s12204-020-2195-y-Figure1-1.png", + "caption": "Fig. 1 Reference frames of an AUV", + "texts": [ + " The main contributions of this paper are listed as follows: a modified error state formulation is introduced to tackle the situation that desired velocities do not satisfy PE condition, which makes it applicable for trajectory tracking control and stabilization simultaneously; a Nussbaum function is used to compensate for the nonlinear term arising from the input saturation. In this paper, we assume that AUVs have axissymmetric appearance and uniform quality, and the impact of roll can be negligible. To study the motion of an AUV, we define an inertial reference frame Oxyz and a body-fixed frame OBxByBzB, as shown in Fig. 1. The mathematic model of AUVs in the 3D space can be written as \u03b7\u0307 = R(\u03b7)\u03bd M\u03bd\u0307 = F (\u03b7, \u03b7\u0307) + Sat(\u03c4 ) + \u03c4w } , (1) with R(\u03b7) =\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 cos\u03c8 cos \u03b8 \u2212 sin\u03c8 cos\u03c8 sin \u03b8 0 0 sin\u03c8 cos \u03b8 cos\u03c8 sin\u03c8 sin \u03b8 0 0 \u2212 sin \u03b8 0 cos \u03b8 0 0 0 0 0 1 0 0 0 0 0 cos\u22121 \u03b8 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 , (2) F (\u03b7, \u03b7\u0307) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 fu fv fw fq fr \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 mvvr \u2212mwwq \u2212 duu \u2212muur \u2212 dvv muuq \u2212 dww (mw \u2212mu)uw \u2212 dqq \u2212WGML sin \u03b8 (mu \u2212mv)uv \u2212 drr \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 , (3) where, \u03b7 = [x y z \u03b8 \u03c8]T; (x, y, z) and (\u03b8, \u03c8) denote the positions (i.e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002483_itec-india.2015.7386862-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002483_itec-india.2015.7386862-Figure14-1.png", + "caption": "Figure 14. Fuel Cell Arrangement - 2", + "texts": [], + "surrounding_texts": [ + "4. Quick position change ability, as jumping out of cover position, stealth mode, and underwater operation without Diesel engine running; 5. Energy saving and regeneration e.g. during braking operation; 6. Electric energy supply to all consumers during standstill of the vehicle, e.g. silent watch; 7. Option for future internal and external supply of high power consumers. 8. According to the wide area of the different tasks greater energy and greater power have to be provided. 9. In addition short duration and long duration of use have to be taken into account. 10. High level of energy is needed for long time application to meet the power requirement of the subsystems e.g. for system initiation/activation or silent watch. High level of power is needed for: The start of the prime mover, Mobility e.g. acceleration, 11. Weapon power supply like ETC-gun, Active armour, Active suspension.\nThe magneto-dynamic storage is a flywheel storage with an integrated electric machine that can be used either as a generator (discharge mode), or as a motor (recharge mode via re-accelerating the flywheel rotor) depending on the momentary needs. The energy carrier of the MDS is a cylindrical rotor made of wound carbon fiber. The rotor\u2019s axle stands on a vertical plane. The motor/generator (M/G) unit is inside the cylindrical rotor, accepting or delivering electric power. To reduce the friction in the bearings most of the rotor\u2019s weight is compensated for by magnetic forces. Air friction is reduced to a minimum as the complete rotor unit runs in a vacuum enclosure.\nThe most important advantages of the MDS compared to other energy storages, i.e. chemical batteries are first of all the high power ability with respect to weight and volume and the indefinite cycle number potential. This is due to the fact that the MDS is an electric machine and is not limited by electrochemical elements. These characteristics open the benefits to use the MDS in military vehicles as an additional energy and power source. The results are the excellent features of the MDS and its related advantages.\nPresently the MDSs reach specific values of energy/mass = 80 MJ/ton and power/mass = 6.5 MW/ton in the laboratory. In the next few years approximately 15 MW/ton will be achievable.\nThis MDS type has been safe proved by the international technical authorities and is authorized for use in public transport applications. With more than 200,000 operation hours and more than 2.5 million kilometres in vehicles this technology stands for an evident reference.\nFuel cells generate electrical energy by an electrochemical reaction (i.e. without combustion). As a result, fuel cells offer a high potential efficiency, and emit exhaust gases comprised solely of water vapour. In contrast to the majority of prime movers, the efficiency of fuel cells is the greatest at partial loads rather than at full load.\nConsequently, other than is the case with internal combustion engines, it is not beneficial to down-size a fuel cell solely on the basis of efficiency considerations \u2013 for example, as a component of a series hybrid system (a relatively small fuel cell and a constant high load, instead of a large fuel cell usually subjected to a partial load). However, since the available fuel cells are still heavy and bulky \u2013 and, above all, expensive \u2013 many projects will nevertheless deploy a small fuel cell. The expected reduction in weight, volume and price will lead to a trend towards the use of fuel cells as a component of the power train. In view of the use of electricity produced by fuel cells to power vehicles the recuperation of the braking energy will continue to be of interest; this can be achieved by the use of a small battery to serve as a buffer, and consequently these systems will continue to be of a hybrid nature. The inherent inertia of reformers may result in the need for Systems employing a reformer to make use of a hybrid power train so as to provide for a rapid response to the operation of the gas pedal. Fuel cells are often regarded as a potential longer-term (several decades) serious alternative to the internal combustion engine.\n1. High potential efficiency, 2. Low emissions, 3. Opportunity to use hydrogen produced from sustainable\nsources, 4. Possibly less maintenance as a result of the absence of\nmoving parts, 5. The opportunity to make use of the existing fuel\ninfrastructure in combination with reformers, 6. Hydrogen is extremely inflammable, and difficult to store, 7. The complete absence of a hydrogen infrastructure for\nsupplies to ordinary vehicles, 8. High costs, 9. The use of a reformer lowers the efficiency, 10. Slow response of the reformer to varying loads, 11. The power train can be of a substantial size,\nSystem integration.\nFuel cells convert chemical energy directly into electrical energy. This process does not involve the Carnot cycle, and consequently extremely high electrical efficiencies (to 80%) are theoretically possible. The fundamental difference between fuel cells and batteries pertains to the fuel and the oxidant. In batteries the fuel and oxidant are stored in the form of a solid or a liquid in the battery (and are regenerated when the battery is charged), whilst fuel cells receive continual supplies of fuel and oxidant in the form of a gas (and sometimes as a liquid) from an external source. Consequently, in analogy with combustion engines, the range of vehicles propelled by fuel cells is determined by the content of the (separate) fuel tank.\n\u2022 ALKALINE FUEL CELLS (AFC) \u2022 SOLID OXIDE FUEL CELLS (SOFC) \u2022 PROTON EXCHANGE MEMBRANE \u2022 FUEL CELLS (PEMFC) \u2022 PHOSPORIC ACID FUEL CELLS (PAFC) \u2022 MOLTEN CARBONATE FUEL CELLS (MCFC)", + "Experiment Tank & Notional Tank - 2 required fuel cells\n\u2022\n\u2022 \u2022\n\u2022\n\u2022\n\u2022\n\u2022\n\u2022\n\u2022\nUse of Fuel Cells have been already incorporated by many research organizations around the globe, which makes it all the more feasible and workable with as some background research material is available. Some more advantages of Fuel Cell Research:\n\u2022 The reduction of the response time from >10 min. to < 1 min; \u2022 The increase of the specific power (W/kg) by means of the\nintegration of the components; \u2022 The reduction of the costs; \u2022 The improvement of the efficiency. \u2022 It is evident that the use of hydrogen as the fuel will require\n(and certainly during the coming ten to twenty years) the deployment of the centralized reforming of (fossil) fuels (for example, as service stations).\n\u2022 This reforming process can be affected in large installations in which issues such as the response time, compactness, etc., are much less critical, and in which higher efficiencies can be much more readily attained.\nENERGY STORAGE" + ] + }, + { + "image_filename": "designv11_34_0003565_j.apm.2020.09.018-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003565_j.apm.2020.09.018-Figure1-1.png", + "caption": "Fig. 1. 3D model of an electromagnetic harmonic movable tooth drive system. 1. flexible ring;2. housing; 3. ion core; 4. coils; 5. air gap; 6. center wheel; 7. movable tooth;8. movable tooth frame (output shaft).", + "texts": [ + " The forced responses of the eccentric flexible ring to exciting current are analyzed. Many useful results are obtained. The results could be used to design and analyze the drive system. 2. Structure and operation principles An electromagnetic harmonic movable tooth drive system composes two main parts: a wave generator and a movabletooth drive. They include eight components: a flexible ring, a housing, a core, coils, an air gap, a center wheel, movable teeth and a movable tooth frame (output shaft). Solid model of the drive system is shown in Fig. 1 . The wave generator comprises a flexible wheel and an electromagnetic core. There is an air gap of length \u03b4 between them. The movable-tooth drive part consists of a center wheel, a movable- tooth frame and various movable teeth. The flexible wheel is a cup-shaped magnetic shell; the bottom of the cup is fixed to the base, and the free end is in contact with the movable teeth. In the movable-tooth drive part, the movable teeth are several balls, and the radial guide grooves with the same number of movable teeth are arranged on the movable-tooth frame, which is used to limit the movement direction of the movable teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001441_978-3-319-05203-8_76-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001441_978-3-319-05203-8_76-Figure3-1.png", + "caption": "Fig. 3 Identification of the angled surfaces", + "texts": [ + " The distribution of the metal powder particles is typically Gaussian with a particle size range from 20 lm (10 % quantile) till 63 lm (90 % quantile). For the processing of aluminum powder a nitrogen inert atmosphere with maximum oxygen content of 0.2 % was used. For the testing of building parameters a three types of test parts were used. The first part of the test was designed to determine the surface roughness at different slope angles to the machine platform. For these purposes a test part was designed with 10 inclination of stepped surfaces (Fig. 3). With such designed surfaces a surface roughness of almost the entire angular range could be obtained using only one test piece. The second test piece was used to determine the dimensional accuracy of the building process. For these purposes a part comprising a hole with a circular (denoted as C) and a square (S) cross-section, the thin wall (T), the radius and angular surfaces has been designed. Geometrical elements were constructed in the range of 0.5\u20135 mm for holes and 0.25\u20132 mm for thin walls", + " Change of this two parameters was separately defined for the outer contour and for the internal volume of the test part. The variation of these parameters within the test series is shown in Tables 1 and 2. To measure the surface roughness a 3D optical profilometer ContourGT-X8 was used. The surface roughness was measured at two points of each angled area. At the place closer the platform (B) and at the place with maximum distance from the platform (T). Areas have been identified depending on the angle between surface normal and normal of the platform, as can be seen on Fig. 3. For determination of the dimensions of the parts and evaluation of geometric accuracy a 3D optical system (ATOS III TripleScan, GOM, Germany) based on active triangulation was used. Before the optical digitization process, it was necessary to use titanium powder for matting the surfaces of the test parts. The maximum thickness of the titanium coating is 6 \u00b1 3 lm. The measured dimensions of the test parts were evaluated in hundredths of a millimeter, therefore additional thickness of matting of the surfaces was neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000185_ijpt.2018.090374-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000185_ijpt.2018.090374-Figure1-1.png", + "caption": "Figure 1 Schematic representation of the planetary wheel hub gears", + "texts": [ + " Planetary hub systems are also particularly compact, yielding highly concentrated tooth meshing contacts. Therefore, a methodical approach capable of predicting the parameters which affect planetary hub gears\u2019 efficiency is the key to achieving efficient systems. The current study presents a parametric analysis of the effect of different extents of tooth longitudinal crowning on the planetary hub gears\u2019 power loss. It also takes into account the influence of roughness of meshing surfaces upon system efficiency. Figure 1 is schematic representation of planetary hub gear. It shows the power flow from the differential gearbox through to the wheel hub. It also shows the transmission ratio, torques and speeds at the different stages of the axle system. Power is transmitted through the sun gear attached to the input shaft. The ring gear is fixed to the transmission housing. The output power is transmitted to the wheels through the carrier. The planetary system comprises three planet gears. It is assumed that there is no misalignment in the planetary system and the input power from the sun gear is equally divided among the planets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003056_5.0002805-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003056_5.0002805-Figure1-1.png", + "caption": "FIGURE 1. Ar.Drone Parrot Bebop and its dimensions", + "texts": [ + " Section 4 describe several simulations that test the mathematical model, along with analyses in the end. Section 5 will close this paper with a few conclusion and suggestion for future works. For this research, an off-the-shelf QUAV named the AR.Drone Parrot Bebop7 was used. This QUAV has the configuration of a body (fuselage) with four symmetrical rotors in X-configuration. Notice that the arms of the rotors, in x and y direction, are different, which means that rolling is easier than pitching. The dimensions of the Bebop are drawn in the vector drawing in Fig. 1, and with other parameters listed in Table 1. Specifications such as weight and performances are listed in the table as well. Based on the configuration and dimension, a typical mathematical model for QUAV can be represented using Simulink blocksets as shown in Fig. 2. From the right, the Simulink model consists of (1) the equation of motion subsystem, (2) the rotor model subsystem, (3) the weight/balance subsystem, and (4) the command subsystem. Each of them represents mathematical equations, elaborated in the following paragraphs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002465_icsima.2014.7047443-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002465_icsima.2014.7047443-Figure2-1.png", + "caption": "Fig. 2. Free-body diagram of the 3-DOF helicopter.", + "texts": [ + " Moreover, few researches focused on decentralized tracking control laws for MIMO models of uncertain plants [8]. The following part of this paper is organized into four sections. In Section II, we briefly describe the Quanser helicopter and its dynamics and mathematic model, followed by a formulation of the robust output regulation for the elevation and pitch axes of the 3-DOF helicopter. In Section III, a new design method of robust output regulator is proposed. Experimental results are presented in Section IV. A 3-DOF bench-top helicopter developed by Quanser shown in Fig. 1 and Fig. 2, where portrayed two DC motors has installed at the ends of a rectangular frame [9]. Those employed to control the 3-DOF lab helicopter. The helicopter frame can move freely to pitch around a long arm and the long arm suspended from a junction who has two degrees of freedom to elevate and travel. These two DC motors are called front motor and back motor separated with independent control signals. The helicopter frame would elevate if both motors were applied with positive voltages, while it would fall with opposite ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001715_978-3-319-19216-1_25-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001715_978-3-319-19216-1_25-Figure1-1.png", + "caption": "Fig. 1. The coordinate system of quadrotor", + "texts": [ + " In this paper the neural network controller for quadrotor steering and stabilizing under the task of flight on path has been presented. It allows to analyse how this type of algorithm can be cooperative with a flying object as a quadrotor. In the first stage of the study, which is described in the paper, the neural network controller was taught on the base of the PID control system. In the second stage of the study the learning process is planned to be extended to fuzzy control system. A quadrotor can be represented as an object with four motors with propellers in cross configuration (Fig. 1). Comparing to a classic helicopter, the division of the main rotor into two pairs of propellers in the opposite direction removes the need of a tail rotor. Usually all engines and propellers are identical, so the quadrotor is a full symmetrical flying object. The detailed mathematical model we based on, can be found in [2]. Although, for clarity the main formulas are presented below. The torque moment around Oy axis: = ( \u2212 ) (1) where b is a thrust coefficient, l is the distance between the propeller's axis and the center of mass of the quadrotor and X1, X2 are rotation speeds of propellers according to the Fig. 1. As the consequence the angle \u0398 called pitch can be observed. The torque moment around Ox axis: = ( \u2212 ) (2) where Y1, Y2 are rotation speeds of propellers. As the consequence the angle \u03a6 called roll can be observed. Quadrotor Navigation Using the PID and Neural Network Controller 267 The join torque around mass center of quadrotor: = ( + \u2212 \u2212 ) (3) where d is so called drag coefficient. As the consequence the angle \u03a8 called yaw can be observed. The above formulas look quite simple and so they are, thus the quadrotor position can be controlled only via propellers rotation speed changes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003605_icuas48674.2020.9213872-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003605_icuas48674.2020.9213872-Figure8-1.png", + "caption": "Fig. 8. Reference flight path of the leader during autonomous filming by a group of UAVs", + "texts": [ + "0 [m/s] SpeedSettings-MaxRotationSpeedCurrent 100.0 [\u25e6/s] 20.0 [\u25e6/s] SpeedSettings-MaxPitchRollRotationSpeedCurrent 300.0 [\u25e6/s] 80.0 [\u25e6/s] PilotingSettings-MaxTiltCurrent 20.0 [\u25e6] 5.0 [\u25e6] PilotingSettings-MaxAltitudeCurrent 30.0 [m] 2.5 [m] PilotingSettings-MaxDistanceValue 2000.0 [m] 3.0 [m] PilotingSettings-NoFlyOverMaxDistance- Shouldnotflyover 0 (false) 1 (true) Gain Gain Gain kP kI kD X-axis 0.69 0.00015 50 Y -axis 0.69 0.00015 50 Z-axis 1.32 0.0003 10.2 \u03b8 default default default \u03c6 default default default \u03c8 0.07 0.00001 0.9 in 3D (see Fig. 8), according to [49], the object model is the smallest possible cuboid in which the object, recorded by the UAV, can be placed. For a constant flight altitude, in the algorithm implemented in the formation controller node [30], it is important to know the vertices of the rectangle O(Xi,Yi), for i = 1, ..., 4, as well as the set value of r. Then, depending on the mutual position of two adjacent corners, it is possible to determine the rotation angle \u03b4 around the Z axis of the global system according to: \u03b4 = arctan (y1 \u2212 y2, x2 \u2212 x1) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003558_icra40945.2020.9197469-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003558_icra40945.2020.9197469-Figure4-1.png", + "caption": "Fig. 4. Drag operation to avoid capture the obstacles: an upward drag step moves the target to an area that contains fewer obstacles ((a) and (b)); an upward push-back step pushes the upper obstacles aside (c) before closing the fingers (d).", + "texts": [ + " During a horizontal zig-zag push operation, the device moves in the xy plane, wherein the resultant vector of the zig-zag motion is equal to Dzigzag and the amplitude ah and number of pushes nhp of the zigzag motion are determined according to the specific picking scenario. For example, the effectiveness of the values may be affected by the peduncle length, fruit weight or the damping ratio of the fruit, which are difficult to calculate. Based on some tests in the farm, in the current system, we tune the ah and nhp to fix values of 20 mm and 5, respectively. C. In-hand Drag Operation Above The Target If an obstacle is located above the target (layer 1), such as the case shown in Fig. 4(a), the gripper may swallow or damage the obstacles when moving upward to capture the target strawberry. Furthermore, the obstacles may stop the fingers closing thus resulting in cutting failure of the target peduncle. To solve this problem, we propose an in-hand drag operation, which is opposite to the push operation as used in other layers. The drag operation allows the gripper to pick the target fruit without capturing unwanted obstacles. The operation comprises an upward drag step to move the target to an area that contains fewer obstacles (Fig. 4(b)) and an upward push-back step that pushes the upper obstacles aside (Fig. 4(c)) before closing the fingers. The push-back step is necessary because when at the drag position (Fig. 4(b)), the peduncle is inclined such that the fruit is difficult to fall due to the static force and easily damaged when the gripper moves up further towards a cutting position. The drag operation is performed only when there are obstacles in the central block CC of the top layer. If the CC is unoccupied, the gripper moves directly upwards to pick the target strawberry. Fig. 5 shows the diagram of the 4959 Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 04,2020 at 03:49:54 UTC from IEEE Xplore. Restrictions apply. calculation method of the drag operation with corresponding to the example in Fig. 4. As shown in Fig. 5(a), to avoid the collision between the gripper and the table, the three blocks LR, CR, RR that are close to the table are skipped for the calculation of the drag direction. Then the drag direction Ddrag in the xy plane can be determined according to the following equation: Ddrag = l\u03a3m 1 Uj/|\u03a3 m 1 Uj | (3) where, Uj is the vector of the jth unoccupied block within the largest group of adjacent unoccupied blocks. The blocks used for calculation are LC, LF, CF, RF, RC. The parameter m is the total number of blocks within the largest group of adjacent unoccupied blocks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003205_s40430-020-02447-7-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003205_s40430-020-02447-7-Figure2-1.png", + "caption": "Fig. 2 Typical geometries of a spiral groove seal", + "texts": [ + "\u00a01, the fundamental elements of spiral groove mechanical seal are the rotor (rotational face) that is rigidly mounted on the shaft, while the stator (stationary face) is spring-loaded and floats axially [21]. The relative rotational motion and normal reciprocating motion between the stator and rotor could generate hydrodynamic pressure, namely the viscous shearing effect and squeeze Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:361 1 3 Page 3 of 11 361 effect. Meanwhile, this phenomenon could be further enhanced by a group of spiral grooves. The geometry of spiral grooves ring is depicted in Fig.\u00a02. Grooves could be manufactured on an arbitrary sealing surface, in this case, in the rotational face [23]. The logarithmic spiral curve is defined as r = rie tan . ri, rg and ro represent the inner radii, groove radii and outer radii, respectively. The pressure at inner radius and outer radius is pi and po. \u03b1 is the spiral angle. \u03c9 represents the angular velocity of rotating ring. The fluid film could be subjected not only to the shear but also the squeeze actions under hydrodynamic loading when the rotational speed and sealed pressure change instantaneously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure3-1.png", + "caption": "FIGURE 3. TEST SETUP FOR STIFFNESS MEASUREMENT AT ATMOSPHERIC PRESSURE FOR VBD BRUSH SEALS.", + "texts": [ + " However, the dependence of brush seal stiffness on the axial pressure drop requires experimentation. Prior to the study of the effect of pressure drop on the seal stiffness, a stiffness test is done on the seal under atmospheric Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/76943/ on 01/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2013 by ASME pressure using a material testing (Instron) machine (Figure 3.) In this test, a segment of the VBD brush seal is mounted on a fixture placed on the Instron machine. A metal shoe with a curvature that matches the seal internal diameter is attached to load cell of the machine. During the test, the shoe undergoes incursion and excursion on the brush seal segment. The reaction force on the shoe from the seal is measure by the load cell, and is plotted as a function of the shoe displacement (for example, see Figure 4.) The slope of the force vs. displacement curve is a measure of the stiffness of the seal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure3.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure3.3-1.png", + "caption": "Fig. 3.3 Pure rotation of a Rigid body over the origin of both base frame and body\u2019s frame", + "texts": [ + " Remark Notice that this transformation is not a linear one since it does not fulfill the superposition principle. Definition 3.1 (The translation director vector \u03bbT ) The translation director vector \u03bbT \u2208 R 3 express the direction of a Cartesian translation d in R 3. If d = \u2016d\u2016 is the magnitude of this translation distance, then the translation director vector is defined as a unit vector \u03bbT d d (3.1) such that d = \u03bbT d 3.2 Rotations For simplicity consider a pure rotation motion were there is no displacement of the origin of the mobile frame w.r.t. the base one as shown in Fig. 3.3. Let 0 be a right-handed reference frame defined with the unit vectors i0, j0 and k0 along respectively to the Cartesian axes x0, y0 and z0 of 0. Let 1 be also a different right-handed reference frame rigidly attached to a body B, defined with unit vectors i1, j1 and k1 along the axes x1, y1 and z1. Any point p in the body has the same physical properties as position, velocity and acceleration regardless wether it is expressed in one frame or another. Let p be the vector representing the position of the point p from the common origin 0; and let 150 3 Rigid Motion p(0) and p(1) be the coordinates expressions of this vector p in frames 0 and 1, respectively as follows: p(0) = p0x i0 + p0y j0 + p0z k0 = \u239b \u239d p0x p0y p0z \u239e \u23a0 \u2208 R 3 p(1) = p1x i1 + p1y j1 + p1z k1 = \u239b \u239d p1x p1y p1z \u239e \u23a0 \u2208 R 3 The relationship that exists between these two vectors can be found by calculating the orthogonal components of each representation along the unit axes of either frame", + " In the figure the first and second rotations are performed about axes of the base frame (extrinsic rotations) while the third is performed with basic rotations about axes on each current frame i.e. while moving (intrinsic rotations). Remark It is remarkable to see that inverting both: the order of rotations and the intrinsic/extrinsic order the final attitude results in the same (see first and third cases). Then for extrinsic rotations (performed relative to the fixed frame) the final rotation matrix is made on by the same basic matrices as for intrinsic rotations (performed relative to current frames) but with reversed order. Recall Fig. 3.3 with all its elements: frames 0, 1 and point p. Let 2 be a third frame, arbitrarily chosen. Then p can also be represented in this frame as p(2). Then there should exist two rotation matrices R2 0 \u2208 SO(3) and R2 1 \u2208 SO(3) which make true the next couple of equations. p(1) = R2 1 p (2) (3.22a) p(0) = R2 0 p (2) (3.22b) Using (3.22a) into the compact form of (3.5), it yields p(0) = R1 0R 2 1 p (2) (3.22c) Notice that in expression (3.22a) the rotationmatrix R2 1 is defined relative to frame 1 and not to the base frame 0", + " Then which set of attitude representation (e.g. any set of Euler angles or quaternions) express the angular velocity with their derivatives? The answer to this question should take in count that the angular velocity vector is unique, whether it is expressed in the inertial frame or body\u2019s frame, regardless of which attitude representation has been chosen. To simplify the analysis of angular velocity, lets study the problem with a pure change in the attitudewithout displacements, i.e. d\u0307 = 0. Take the example of Fig. 3.3, where both frames has the same origin and suppose that the body B is moving with a angular velocity \u03c9 \u2208 R 3, and let R(t) = R(\u03b8(t)) \u2208 SO(3) be the instantaneous rotation matrix from frame 0 to frame 1 at an instant t , so that the position of point p at that instant is d(0) p = R(t)d(1) p (3.41) Then the instantaneous linear velocity of point p is given by the time derivative of last expression as 3.3 The Rigid Motion Kinematics 167 d\u0307 (0) p = R\u0307(t)d(1) p + R(t)d\u0307 (1) p = R\u0307RT d(0) p (3.42) where d\u0307 (1) p = 0 because it is defined relative to the frame attached to the body, hence constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003272_j.promfg.2020.05.027-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003272_j.promfg.2020.05.027-Figure2-1.png", + "caption": "Figure 2. Illustration of the working principle of waveguide-based L&T vibrator.", + "texts": [ + " Compared to the first method of using straight slits arrays, the helical grooves are more difficult to be machined due to the complex geometric shapes, even though the torsional conversion efficiency sometimes can be higher than that of the diagonal slits. However, both the two methods require the resonant frequency of the torsional vibration mode to be the same as that of their longitudinal mode. This is challenging particularly when they are sharing only one node plane. Inspired by the idea of modifying the structure of the longitudinal-excited vibrator, this paper proposed to design a waveguide structure with internal and external geometric features. An illustrative example of L&T vibrator design is shown in Figure 2. An acoustic waveguide structure can guide a wave propagation direction with minimal loss of energy, hence high vibrations conversion efficiency. Therefore, a waveguide can constrain the wave propagation along the defined path to further achieve a 2-dimensional hybrid vibration. Our objective is that the overall composite vibration wave can form a hybrid L&T vibration on the same plane, as shown in Figure 2. As shown in Figure 2, the new waveguide-based vibrator was designed by using a CAD software SolidWorks 2019. The basic dimension of the vibrator is \u03a628 mm \u00d737 mm, as shown in Figure 3. The diameter of the flange is \u03a628 mm. Based on the size of the flange, the wave propagation path needs to be analyzed and defined. To achieve a hybrid L&T vibration, a circular array of helical paths is used in which each helical path can guide the wave to a certain angle (see Figures 2 and 3). The defined wave path is a circular helix trajectory, which can be represented as follows: { \ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03) = \ud835\udc37\ud835\udc37 2 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udf03\ud835\udf03 \ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03) = \ud835\udc37\ud835\udc37 2 \ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udf03\ud835\udf03 \ud835\udc67\ud835\udc67 = \ud835\udc3f\ud835\udc3f 2\ud835\udf0b\ud835\udf0b \ud835\udf03\ud835\udf03 , \ud835\udf03\ud835\udf03\ud835\udf03\ud835\udf03[0, \ud835\udf0b\ud835\udf0b] (1) Where D is the diameter of the circular helix, and the L is the length of the rods. Based on the wave path, six patterned helical rods were created as the final waveguide structure (see Figure 2). The diameter of the rods was designed as \u03a64.5 mm, which is determined by the maximum sustainable stress based on our FE analysis, as discussed in the next section. The design and configuration of the assembled new L&T vibrator are shown in Figure 4. To obtain high electromechanical conversion efficiency and reasonably stable performance, a bolt-clamped type Langevin transducer is used to generate the longitudinal vibration. The Lead Zirconate Titanate (PZT-8) piezoelectric ceramic rings are polarized along the thickness direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002481_urai.2016.7625728-FigureI-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002481_urai.2016.7625728-FigureI-1.png", + "caption": "Fig. I. Concept of control signal broadcast. Sending control signals through a laptop.", + "texts": [], + "surrounding_texts": [ + "The technique is simple to use for the formation control in this study. Four drones will be moving the same for the same control signal from the laptop. And each of the drones share the position information of the GPS mounted on the drone in a mesh network of ZigBee. Each of the drones calculates how much to move through a shared position information. Before take-off, each of the drones calculates the distance between each drone from own position information and position information of the other drones. And each of the drones maintains a calculated distance during the flight. The initial position of the drone will be placed in the desired formation. After, each of the drones flies according to the control signals coming from the laptop. Fig. 1 shows that the control signal transmission. A communication module connected to a laptop can be seen that sending a control signal to each of the drones. 3. Implementation 3.1 Quadrotor specification The flight control unit (FCU) has used to implement is the Pixhawk of the 3 D Robotics, which is an outstanding company in this field. By modifying an open source offering the 3D Robotics, we used for FCU. And the other part of the drone, frames, brushless direct current (BLD C) motors, motor drivers and batteries, were purchase . d separately. Each drone was produced in the all same specI- fication. The size of the drone used in the experiments is 500mm. Also the size of a propeller is I I-inch. In general case, it is difficult to fix the drone in the air. The drone is flowing because of less friction in the air. Due to various reasons, these problems appear. Among them, the biggest reason is the location of the center of gravity of the drones. To position the center of gravity of the drones in the center is very difficult. So, in order to prevent that the drones flow, we correct the position by using GPS. As mentioned earlier, however, GPS has a positioning error of about several meters. For this reason, to reduce the positioning error, we used an optical camera with GPS. Optical camera knows own moving direction for finding the pattern of the floor. Position control mode using the optical camera and GPS so that makes drones only move for control signals. Typical current drones are only one drone operations thorugh a single transmitter. In order for the formation co- ntrol of 4 drones in the same direction at the same time, it must be connected to the four drones in one transmitter. For this implementation, we communicate I-to-4 communication using ZigBee. At this time, baud rate of ZigBee communication use 115600 bit per second and distance of laptop is within a radius of 200 meters. When master ZigBee connected to the laptop broadcasts the control signal to others, four slave Zig Bee connected to each of the drones catch the control signal and move to match the control signal. In order to control through a laptop for each of the drones, we should know exactly what the signal from the transmitter to send quadrotor. By analyzing the code in an open source offering in the company, you can see what the signal is going through the transmitter. Through this method we can know some important signals, which are arming signal, throttle, roll, pitch and yaw. After confirmation of these signals, it created a console program that can continue to send some signal previously confirmed on the laptop. And through the modified open source only the signal, which is not the transmitter signal, coming through the ZigBee connected to the drone is to make the movable quadrotor. 3.3 The additional equipment in system Fig. 3 is a ground control system provided by pixhawk.org installed on the laptop. Using this software program, we can monitor the value of each variable and the current state of the drones. In addition, the GPS position imformation is also displayed on the program. As previously mentioned in this study, the GPS module was used to estimate the position of a quadrotor in outdoor environment. And an optical camera was used to reduce the effect of a positioning error of the GPS module. Fig. 4 is a system model showing the entire system. It shows the correlation of the respective device." + ] + }, + { + "image_filename": "designv11_34_0002759_2013.17855-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002759_2013.17855-Figure6-1.png", + "caption": "Figure 6. Vehicle Inclinations", + "texts": [ + " And is a yaw angle of the vehicle. Actually, vector C is constant, and vector is given as a positioning data of GPS, and also is given by the IMU. A global map can be obtained by temporal integration i.e. integrate range data based on global coordinates. Then, this map represents whole surrounding scene, as shown in Figure 5. C i\u03b8 P i\u03b8 n This system aims to apply for off-road environment. Therefore, vehicle inclinations were also considered. Vehicle inclinations of roll and pitch angle were defined as shown in Figure 6. Eqn. (3) is built by added rotational transformation term to Eqn. (2). ( ) PClRRTRw i\u03b1\u03b2\u03b8i +\u2212= (3) \u2212= \u03b1\u03b1 \u03b1\u03b1 cossin0 sincos0 001 \u03b1R Digital Elevation Map (DEM) A global map based on point data is not appropriate for description of 3D environment. In the study, Digital Elevation Map (DEM) was applied for representation. On DEM, the environment is divided into uniform grids. The highest point data in a grid was adapted as height data. Using this height, DEM describes three-dimensional environment (Yoshimitsu, 2000)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003229_robosoft48309.2020.9115971-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003229_robosoft48309.2020.9115971-Figure1-1.png", + "caption": "Fig. 1. Structure of the phantom showing the silicone layers and the sensor layer and the line sensor morphology used.", + "texts": [ + " This relationship has been previously verified [20]. The sample phantom has been developed by casting layers of EcoFlex 00-10 sandwiched together with sensing layers of CTPE. The height of the bottom silicone layer determines the range of possible deformation and compression of the phantom. This is followed by a sensing layer. The sensing layer consists of a x-direction sensing layer, a thin layer of silicone and then the y-direction sensing layer, sandwiched together using Dow-Corning Silicone adhesive (Fig. 1). Each sensor layer consists of an array of CTPE sensors, each straight lines spanning the length of the phantom, separated by the distances dx and dy in the two respective directions. Each layer enables the respective localization of the deformation in the x and y direction. For n sensors in each of the x and y directions, each with a sensor length, l, the sensor morphology can be described by Mnx = [nd, l],Mny = [l, nd]. The change in resistance in each sensor reflects the change in length caused by the deformation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure7-1.png", + "caption": "Fig. 7. Temperature in the wave cooling duct for different flow rates (2.5, 5, 10 l/min).", + "texts": [ + " The 2- way coupled model (Fluent) is used in Fig. 6a, whereas the non-coupled model (COMSOL) is used in Fig. 6b as a comparison. For the latter in Fig. 6b, the temperature along a radial line that starts in the stator core between the slots and ends in the coolant is shown. There is a 3 to 6 degrees difference between the aluminum and the coolant, representing a thermal resistance. The temperature distribution in the wave cooling duct for different flow rates (2.5, 5, 10 l/min), when using the 2-way coupled model is shown in Fig. 7 (for the same loss value). It can be seen that the temperature is rather evenly distributed for the high flow rate of 10 l/min whereas there is a more distinct difference between outlet and inlet temperature for the low flow rate. 869 Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 15,2021 at 16:48:44 UTC from IEEE Xplore. Restrictions apply. The resulting temperatures achieved when having a spiral cooling duct is quite like that of the wave cooling duct, as seen in the comparison in Table IV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003203_j.est.2020.101417-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003203_j.est.2020.101417-Figure4-1.png", + "caption": "Fig. 4. FEM model for the numerical calculations.", + "texts": [ + " Following the modal analysis, a harmonic sine sweep was employed between 5 and 100 Hz, the steps of which are presented in the following section. Lastly, a random vibration profile was applied to the battery block and the response of the structure was obtained. The representative battery pack considered in the study was divided into four main components: the fixture, the bottom plate, the battery cells and the connection elements and the top plate, as shown in Fig. 3. The assembled FEM model of the battery block and the fixture are illustrated in Fig. 4. In the analysis, linear elastic material behaviour was assumed, and the small deflection theory was used. In addition, no nonlinearities were included. The damping coefficients employed in the related simulations were selected based on a previous study [11] by the authors. The FEM model included a total of approximately 15,918 elements and 44,412 nodes. The material properties pertaining to the battery block used in the FEM analysis are given in Table 2. The sine sweep vibration profiles in all directions result in almost similar outputs for the circular, 6 \u00d7 6, 4 \u00d7 9 and 3 \u00d7 12 topologies as shown in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000318_j.finel.2015.07.008-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000318_j.finel.2015.07.008-Figure1-1.png", + "caption": "Fig. 1. A geared rotor system composed of five shafts.", + "texts": [ + " In this paper, the dynamics of an integrally centrifugal compressor with five parallel shafts is investigated, where the changing parameters of TPJBs are considered as the main causes of the severe vibrations mentioned above. Admittedly, the \u201ccritical load\u201d may not be caused only by the changing parameters of TPJB, but it is certain that it is one of the most important reasons. Based on the finite element (FE) model of the rotor system, the response reproduced the vibration phenomenon in the starting-up process. These results will provide guides for the integrally centrifugal compressor design. compressor with five parallel shafts A schematic of the geared parallel-rotor system is shown in Fig. 1, which is a prototype of an integrally geared centrifugal compressor with 5 shafts coupled by helical gears in parallel arrangement. Each shaft is supported by two TPJBs. The test rig is driven by an AC motor connected to the input shaft (I). The dummy impellers or discs are fixed on the three output shafts (O1, O2, O3). These shafts are all connected to a big main gear (M) by different gears. The transmission ratios of the three output shafts are 2.72 (I to O1), 3.25 (I to O2) and 5.10 (I to O3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002133_j.ifacol.2015.11.197-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002133_j.ifacol.2015.11.197-Figure2-1.png", + "caption": "Fig. 2. Model of the acceleration controlled cart-pendulum.", + "texts": [ + " We take advantage of the low state dimension of this example for graphical analysis. Included control and energy phase portraits (Figs. 3 and 5) illustrate how the proposed switching structure yields SAC controllers that leverage free dynamics whether or not manifold tracking goals are included in costs. The example also shows how trajectories evolve through phase space and onto stable manifolds under these different objectives. This section pertains to the frictionless, acceleration controlled cart-pendulum in Fig. 2, with length r = 2 m, mass m = 1 kg, and gravity g = 9.81 m s2 . The uncontrolled pendulum is a Hamiltonian system with energy E(\u03b8, \u03b8\u0307) = 1 2 mr2\u03b8\u03072 +mgr(cos \u03b8 + 1) such that the free dynamics are ( \u03b8\u0307 \u03b8\u0308 ) = f1(\u03b8, \u03b8\u0307) = ( \u03b8\u0307 g r sin(\u03b8) ) . The dynamics of the acceleration controlled cart are defined as x\u0308c(t) = u(t), so that the controlled mode 2 is f2(\u03b8, \u03b8\u0307, u \u2217 2) = ( \u03b8\u0307 g r sin(\u03b8) + u\u2217 2 r cos(\u03b8) ) . While the pendulum\u2019s downward equilibrium, (\u03b8\u0304, \u02d9\u0304\u03b8) = (\u03c0, 0), is stable, the upright equilibrium, x\u0304 := (\u03b8\u0304, \u02d9\u0304\u03b8) = (0, 0), is not", + " for (\u03b8, \u03b8\u0307) \u2208 W s loc(x\u0304) \u222a Wu loc(x\u0304) it holds E(\u03b8, \u03b8\u0307) = E(x\u0304) = 2mgr and we define E\u0304 := E(x\u0304). Locally around x\u0304, the stable manifold is given by W s loc(x\u0304) = { (\u03b8, \u03b8\u0307) \u2223\u2223\u2223\u2223 \u03b8\u0307 = \u2212sign(\u03b8) \u221a 2 g r (1\u2212 cos \u03b8) } and the unstable manifold by the same relation with opposite sign. Globally, the stable and unstable manifolds form a so called homoclinic orbit (cf. red curve in Fig. 3). 2015 IFAC ADHS October 14-16, 2015. Atlanta, USA Alex Ansari et al. / IFAC-PapersOnLine 48-27 (2015) 335\u2013342 339 Fig. 2. Model of the acceleration controlled cart-pendulum. 4. ENERGY TRACKING FOR THE CART-PENDULUM Demonstrating a scenario where stable manifold tracking reduces to energy tracking, this section includes swing- up results for a cart-pendulum. We take advantage of the low state dimension of this example for graphical analysis. Included control and energy phase portraits (Figs. 3 and 5) illustrate how the proposed switching structure yields SAC controllers that leverage free dynamics whether or not manifold tracking goals are included in costs. The example also shows how trajectories evolve through phase space and onto stable manifolds under these different objectives. This section pertains to the frictionless, acceleration controlled cart-pendulum in Fig. 2, with length r = 2 m, mass m = 1 kg, and gravity g = 9.81 m s2 . The uncontrolled pendulum is a Hamiltonian system with energy E(\u03b8, \u03b8\u0307) = 1 2 mr2\u03b8\u03072 +mgr(cos \u03b8 + 1) such that the free dynamics are ( \u03b8\u0307 \u03b8\u0308 ) = f1(\u03b8, \u03b8\u0307) = ( \u03b8\u0307 g r sin(\u03b8) ) . The dynamics of the acceleration controlled cart are defined as x\u0308c(t) = u(t), so that the controlled mode 2 is f2(\u03b8, \u03b8\u0307, u \u2217 2) = ( \u03b8\u0307 g r sin(\u03b8) + u\u2217 2 r cos(\u03b8) ) . 4.1 Stable manifold and cost formulation While the pendulum\u2019s downward equilibrium, (\u03b8\u0304, \u02d9\u0304\u03b8) = (\u03c0, 0), is stable, the upright equilibrium, x\u0304 := (\u03b8\u0304, \u02d9\u0304\u03b8) = (0, 0), is not" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure6.9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure6.9-1.png", + "caption": "Figure 6.9 A fork-on-surface joint.", + "texts": [], + "surrounding_texts": [ + "Joints and Their Kinematic Characteristics 133\nA point-on-plane joint is illustrated in Figure 6.7. It is so called because the kinematic elements ab and ba are in contact with each other at a single point Oba. It may also be called a point-plane contact joint. With respect to ab, Oba moves in the plane formed by u\u20d7(ab)\ni and u\u20d7(ab) j . Meanwhile, ba rotates as if there is a spherical joint centered at Oba. Therefore, the mobility of a point-on-plane joint is \ud835\udf07ab = 5 and it is characterized by the following orientation and location equations, which involve a set of five joint variables such as {\ud835\udf19ab, \ud835\udf03ab, \ud835\udf13ab, xab, yab}.\nC\u0302(ab,ba) = en\u0303ab1\ud835\udf19ab en\u0303ab2\ud835\udf03ab en\u0303ab3\ud835\udf13ab (6.29) r(ab)\nab,ba = uixab + ujyab (6.30)\nThe unit column matrices nab1, nab2, and nab3 that take place in Eq. (6.29) are selected according to similar considerations explained in Section 6.2.3 for a spherical joint.\nA point-on-surface joint is illustrated in Figure 6.8. It is so called because ab and ba are in contact with each other at a single point Oba and ab has the shape of a specified surface ab. The reference frame ab is to be selected in such a way that the function f ab(x, y, z) that describes the surface ab has the simplest expression. For example, if the", + "134 Kinematics of General Spatial Mechanical Systems\nsurface is ellipsoidal, then Oab must be placed at the center of the ellipsoid and the basis vectors of ab must be aligned with the principal axes of the ellipsoid. With respect to ab, the contact point Oba moves so as to remain on the surface ab. Meanwhile, ba rotates as if there is a spherical joint centered at Oba. Therefore, the mobility of this joint is \ud835\udf07ab = 5 even though it is represented by the following set of six joint variables.\n{xab, yab, zab, \ud835\udf19ab, \ud835\udf03ab, \ud835\udf13ab} (6.31) This joint is characterized by the following orientation and location equations together with the surface equation, which constitutes a constraint among the three locational joint variables.\nC\u0302(ab,ba) = en\u0303ab1\ud835\udf19ab en\u0303ab2\ud835\udf03ab en\u0303ab3\ud835\udf13ab (6.32) r(ab)\nab,ba = uixab + ujyab + ukzab (6.33) fab(xab, yab, zab) = 0 (6.34)\nThe unit column matrices nab1, nab2, and nab3 that take place in Eq. (6.32) are selected according to the similar considerations explained in Section 6.2.3 for a spherical joint.\nAs a typical surface example, an ellipsoidal surface is described by the following equation.\nfab(xab, yab, zab) = (xab\u2215aab)2 + (yab\u2215bab)2 + (zab\u2215cab)2 \u2212 1 = 0 (6.35) In Eq. (6.35), aab, bab, and cab are the semi-axis lengths of the ellipsoidal surface.\nba are connected to each other with two point-on-surface joints. The contact points", + "Joints and Their Kinematic Characteristics 135\non ab and ba are denoted as {Pab, Qab} and {Pba, Qba}. The reference frame ab is to be selected in such a way that the function f ab(x, y, z) that describes the surface ab has the simplest expression. As for the reference frame ba, it is selected here so that u\u20d7(ba)\ni is directed from Qba to Pba and u\u20d7(ba)\nk is aligned with the handle of the fork, i.e. with the line that is normal to the line segment between Qba to Pba at its midpoint Oba. The contact points on ba are located by the position vectors r\u20d7bap and r\u20d7baq with respect to Oba. Let the distance between the prongs of the fork be 2dba. Then, r\u20d7bap and r\u20d7baq are represented by the following column matrices in ba.\nr(ba) bap = + dbaui r(ba) baq = \u2212 dbaui\n} (6.36)\nThe relative position of ba with respect to ab is represented by six joint variables, which are the coordinates of Oba (i.e. xab, yab, zab) and the Euler angles (i.e. \ud835\udf13ab, \ud835\udf03ab, \ud835\udf19ab) of an appropriately selected sequence. However, these six variables are subject to two constraint equations, which state that the coordinates of the contact points must satisfy the surface equation specified for ab. Therefore, the mobility of this joint is \ud835\udf07ab = 4. Its characteristic equations are written as follows by expressing the relative orientation matrix with the k\u2013j\u2013i sequence, which will be seen to be an appropriate selection later in Eqs. (6.41) and (6.42).\nC\u0302(ab,ba) = eu\u0303k\ud835\udf13ab eu\u0303j\ud835\udf03ab eu\u0303i\ud835\udf19ab (6.37) r(ab)\nab,ba = uixab + ujyab + ukzab (6.38)\nfab(xabp, yabp, zabp) = 0 fab(xabq, yabq, zabq) = 0\n} (6.39)\nIn Eq. Set (6.39), the arguments of the surface function f ab(x, y, z) are the i, j, k components of the position vectors r\u20d7abp and r\u20d7abq, which locate the companion contact points Pab and Qab of ab with respect to Oab. These vectors are represented by the following column matrices in ab.\nr(ab) abp = r(ab) ab,ba + C\u0302(ab,ba)r(ba) bap r(ab) abq = r(ab) ab,ba + C\u0302(ab,ba)r(ba) baq\n} (6.40)\nWhen Eqs. (6.36)\u2013(6.38) are substituted, Eq. Set (6.40) provides the following detailed equations.\nr(ab) abp = uixab + ujyab + ukzab + dbaeu\u0303k\ud835\udf13ab eu\u0303j\ud835\udf03ab ui (6.41)\nr(ab) abq = uixab + ujyab + ukzab \u2212 dbaeu\u0303k\ud835\udf13ab eu\u0303j\ud835\udf03ab ui (6.42)\nEquations (6.41) and (6.42) show that the selected k \u2212 j\u2212 i sequence is appropriate, because it has provided a simplification owing to the fact that eu\u0303i\ud835\udf19ab ui = ui. Incidentally, considering this simplification criterion, the j\u2212 k \u2212 i sequence could also be selected as another appropriate sequence, if desired.\nNote that u\u0303iuj = uk , u\u0303juk = ui, and u\u0303kui = uj according to the right-handedness property. Therefore, referring to Section 2.7 of Chapter 2, it can be shown that\neu\u0303i\ud835\udf03uj = ujc\ud835\udf03 + uks\ud835\udf03 ut j eu\u0303i\ud835\udf03 = ut j c\ud835\udf03 \u2212 ut ks\ud835\udf03\n} (6.43)" + ] + }, + { + "image_filename": "designv11_34_0001076_cjme.2013.01.011-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001076_cjme.2013.01.011-Figure1-1.png", + "caption": "Fig. 1. FPMLSM structure", + "texts": [ + " BRANDENBURG, et al[5], BR\u00dcCKL[6], GROLLING, et al[7], and ROCHE, et al[8], used current oversampling technology to suppress noise in current signal. KIM, et al[9] and SU, et al[10], used special velocity estimator to minimize speed noise. YUSUKE, et al[11], used 2-degreesof-freedom control structure to suppress the standstill vibration of linear motor. Through these work, the linear motor\u2019s vibration was significantly suppressed. However, these studies did not pay much attention to motor\u2019s normal vibration. FPMLSM\u2019s vibration includes longitudinal direction (X-direction) and normal direction (Z-direction) as shown in Fig. 1. Generally, the X-direction vibration attracts more research interests than the Z-direction vibration, because MENG Fanwei, et al: Adaptive PI Control Strategy for Flat Permanent Magnet Linear Synchronous Motor Vibration Suppression \u00b712\u00b7 the X-direction is motion direction and directly relates to position accuracy. Actually, the FPMLSM\u2019s Z-direction stiffness is less than X-direction for the mover\u2019s flat shape. As a result, FPMLSM will be easy to deform and vibrate along the Z-direction. Further more, the area projected along the Z-direction is larger than the area projected along the X-direction, so more acoustic noise will be generated along the Z-direction even when they have the same vibration amplitude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003272_j.promfg.2020.05.027-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003272_j.promfg.2020.05.027-Figure3-1.png", + "caption": "Figure 3. Design of the waveguide-based L&T vibrator with major dimensions.", + "texts": [ + " An acoustic waveguide structure can guide a wave propagation direction with minimal loss of energy, hence high vibrations conversion efficiency. Therefore, a waveguide can constrain the wave propagation along the defined path to further achieve a 2-dimensional hybrid vibration. Our objective is that the overall composite vibration wave can form a hybrid L&T vibration on the same plane, as shown in Figure 2. As shown in Figure 2, the new waveguide-based vibrator was designed by using a CAD software SolidWorks 2019. The basic dimension of the vibrator is \u03a628 mm \u00d737 mm, as shown in Figure 3. The diameter of the flange is \u03a628 mm. Based on the size of the flange, the wave propagation path needs to be analyzed and defined. To achieve a hybrid L&T vibration, a circular array of helical paths is used in which each helical path can guide the wave to a certain angle (see Figures 2 and 3). The defined wave path is a circular helix trajectory, which can be represented as follows: { \ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03) = \ud835\udc37\ud835\udc37 2 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udf03\ud835\udf03 \ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03) = \ud835\udc37\ud835\udc37 2 \ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udf03\ud835\udf03 \ud835\udc67\ud835\udc67 = \ud835\udc3f\ud835\udc3f 2\ud835\udf0b\ud835\udf0b \ud835\udf03\ud835\udf03 , \ud835\udf03\ud835\udf03\ud835\udf03\ud835\udf03[0, \ud835\udf0b\ud835\udf0b] (1) Where D is the diameter of the circular helix, and the L is the length of the rods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000422_transducers.2015.7180945-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000422_transducers.2015.7180945-Figure2-1.png", + "caption": "Figure 2: Schematic cross-section of the adaptive lens.", + "texts": [ + " When taking into account the material stiffness, this profile will be rather hyperbolic as the piezo will resist the bending. One can expect that a combination of both modes of operation creates a combined boundary condition and gives thus two degrees of freedom to control the aspherical behavior of the membrane. Still, we so far only described a membrane. To create a lens, we need to add a refractive medium below the membrane. In [7], this was done with a fluid, but now, we will join the membrane to an elastomer lens body as shown in fig. 2. Because of the counter-pressure upon deformation, the thickness and elasticity of this body will also influence the tuning range. 978-1-4799-8955-3/15/$31.00 \u00a92015 IEEE 399 Transducers 2015, Anchorage, Alaska, USA, June 21-25, 2015 The fabrication is a straightforward pick-and-place process (fig. 3): First the piezo and the glass are structured by laser ablation and glued together using a rigid polyurethane. The elastomer body is molded directly onto the lens substrate in a milled polymer mold" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003713_iecon43393.2020.9254216-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003713_iecon43393.2020.9254216-Figure2-1.png", + "caption": "Fig. 2. Thumb mechanism axes articulation.", + "texts": [ + " Each element has been cooked at 100\u00ba during 1 hour in an induction oven after being cleaned. This process increased PLA resistance and hardness. The cartilage between each phalanges is done by transparent nail porcelain. In our case the human hand scanned is a hand of a person with high characteristics that entails a big hand dimensions. Each finger consist by 3 phalanges except the thumb that have 2 phalanges and a metacarpal bone. The carpal bone is designed to be fixed with the metacarpal bones for the 4 fingers Fig.2. Each finger has 3DoF there are 15 DoF for the whole hand. The 4 fingers have 12 DoF that realize the flexion/extension and the thumb realizes 2DoF of flexion/extension and 1 DoF for abduction/adduction. The abduction/adduction of the thumb is realized with the metacarpal bone and carpal bones. The abduction/adduction for the join on MPC and CMC has been ignored, that reduce fingers movement only to flexion/extension Fig.2. The abduction/adduction for the join of the Dip and PIP has been ignored. Afterwards, the parts were assembled one by one. 601 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 20,2021 at 06:34:25 UTC from IEEE Xplore. Restrictions apply. The prosthetic hand contains all joins and ligaments as the real human hand. The ligaments and tendons are produced with rubber adapted to each finger join individually. The volar plate and collateral ligaments are developed by rubber with a hardness of 7kg/m and thickness of 0,7mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001506_s00170-014-6017-y-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001506_s00170-014-6017-y-Figure4-1.png", + "caption": "Fig. 4 The gear\u2013rack hardening layer", + "texts": [ + " Twomaterial properties of the gear and rack must be defined in the engineering data since the ANSYS Workbench does not have their properties and constitutive relations. Yield stress, tangent modulus, elastic modulus, and Poisson\u2019s ratio of the gear and rack are input according to Table 2 so as to get the BKIN constitutive curve of gear\u2013 rack hardening layer, as shown in Fig. 3. The surface of the gear and rack can be hardened to improve the surface strength of tooth. Treatment layer thickness is about 3 mm. Figure 4 displays the gear\u2013 rack model treated and the indicating region with arrows is the hardening layer. This paper considers the contact elastic\u2013plastic finite element analysis of the gear\u2013rack contact hardening layer and core un-hardening layer. The ANSYS Workbench provides a tool of \u201cBody Split and Form New Part\u201d for adding different materials to different parts of the gear and rack. These tools are designed to divide a part into different regions according to the requirement to add different materials in different areas of the same parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000340_jae-141878-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000340_jae-141878-Figure10-1.png", + "caption": "Fig. 10. Deflection for 76.2 mm model (a) FEA with just the inner and cage rotor, (b) radial deflection.", + "texts": [ + " The magnetomechanical iterative calculation is shown in Fig. 7(c), it takes 5 iterations to converge. A summary of the deflection results for larger axial lengths is shown in Table 3. The deflection analysis presented above has been validated by experimentally measuring the deflection when using an MDL400 Tesa-Hite height gauge for the case when only the inner rotor and one cage rotor bar is present. Figure 9 shows the inner and outer rotor of the FFMG with 76.2 mm active stack length. The deflection obtained from FEA is shown in Fig. 10. The comparison with experimental deflection is shown in Fig. 11. The deflection accuracy is not as high as for the case when using a 6 inch axial length [6] however the results are sufficiently accurate for design purposes. In this paper it has been shown that the amount of radial deflection on the magnetic gear cage rotor bars for both the scaled-up and subscale magnetic gear increases linearly with axial length. The level of deflection is not significantly changed by the cage rotor diameter. The paper also demonstrates that, for the axial lengths considered, the flux-focusing magnetic gear torque density is not significantly affected by the changes in the axial length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000043_j.prostr.2020.02.048-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000043_j.prostr.2020.02.048-Figure1-1.png", + "caption": "Fig. 1. Projected samples location on the building platform: a) schematic representation of planned subdivision , b) appearance of the job at the end of the process.", + "texts": [ + " The building job involved the production of samples with their main axes along different building directions and a reasonable number of specimens have been produced with following geometry: metallographic samples with dimensions 25 x 15 x 6 mm (length x depth x height) and standard tensile specimens with nominal diameter 5.0 mm (according to ISO 6892-1:2016 Standard). They have been homogeneously located on the building platform, assuming to subdivide the working area in 9 (nine) quadrants, as schematized in Fig. 1(a). As-built samples, still attached to the building platform, can be seen in Fig.1(b). Metallographic samples have been classified in 6 (six) different groups, depending on the condition under which they were tested. Specific heat treatment conditions (six in total) are reported in Table 2. Metallographic investigations and hardness indentation tests (Rockwell C and Vickers) have been performed on properly pre-processed samples, as per specific identified Standard, on both as-built (0 group) and heat treated (A1, A2, B1, B2, C groups) specimens. HRC and HV1 hardness tests have been performed on the same specimens and compared for cross validation; HV1 hardness tests were also performed to support metallurgical investigations and the interpretation of specific microstructure features" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure21-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure21-1.png", + "caption": "Figure 21: Unpressurized nonrotating rotor interference CAE analysis with C3D8I elements for test seals \u2013 VM stress profile at 0.6mm rotor interference", + "texts": [ + "org/about-asme/terms-of-use The displacement contour in the unpressurized-nonrotating rotor interference simulation with C3D8I elements is detailed in Figure 20 for 0.6 mm radial interference. Due to frictional effects and a cant angle of 45o, the rotor interference deflects the bristles both in the radial and tangential directions. The magnitude of the maximum bristle tip displacement is around 0.832 mm for a rotor interference of 0.6 mm. Usage of C3D8I elements instead of B31 elements enables better viewing of displacement profile with smoother contour. The VM stress profile at 0.6 mm rotor interference is given in Figure 21 for FE model using C3D8I elements. The smooth stress contour of C3D8I elements gives a better understanding of the stress profile of the bristles at unpressurized rotor interference conditions. The maximum VM stress is observed at the pinch point, which is direct result of loading condition of an unpressurized seal. The bristles act like a cantilever beam constrained from top and transversely loaded from the bristle tip. The maximum VM stress level at the pinch point is equal to 115.2MPa for rotor interference level of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001636_1.4032400-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001636_1.4032400-Figure1-1.png", + "caption": "Fig. 1 Relative positions of the teeth-skipped worm and the gear", + "texts": [ + " Thus, the analytical results such as profile deviations or pitch deviations cannot be derived from the conventional single-flank composite testing. To solve this problem, Huang Tongnian et al. improved the principle of the single-flank composite testing in 1969 and the proposed method is called GIE technology today [15,21,22] (Huang Tongnian wrote [15] under the pseudonym of Huang Lili). The main improvement is that the number of pairs of contacted flanks during GIE measuring is never greater than one by adopting a new type of masters, the so-called \u201cteeth-skipped worm.\u201d Figure 1 shows the relative positions of the teeth-skipped worm and the gear to be measured during inspection. Two highresolution encoders are mounted on the axes of the worm and the gear, respectively. The GIE curves can be acquired by the data from the encoders. In order to investigate the meshing process between the teethskipped worm and the gear, it is assumed that there is only one tooth on the worm and only one tooth on the gear; thus, the full process of meshing is illustrated as shown in Fig. 2. As shown in lower part of Fig", + " 5 /AOwC q; h\u00f0 \u00de \u00bc h\u00fe cos rb1=q\u00f0 \u00de (7) CD q; h\u00f0 \u00de \u00bc /AOwC q; h\u00f0 \u00de rb1 kb1 (8) zM q; h\u00f0 \u00de \u00bc CD q; h\u00f0 \u00de rb1 tan arccos rb1=q\u00f0 \u00de\u00f0 \u00de tan kb1\u00f0 \u00de (9) For the purpose of simplifying the calculation, the origin of the Zw-axis can be set to the pitch point determined by the radius of the pitch circle zp \u00bc zM rw1; 0\u00f0 \u00de (10) where rw1 denotes the radius of the pitch circle of the worm. Then, the equations representing the worm tooth surface are described as xM q;h\u00f0 \u00de yM q;h\u00f0 \u00de zM q;h\u00f0 \u00de 0 BB@ 1 CCA\u00bc qcos h\u00f0 \u00de qsin h\u00f0 \u00de CD q;h\u00f0 \u00de rb1 tan arccos rb1=q\u00f0 \u00de\u00f0 \u00de tan kb1\u00f0 \u00de zp 0 BB@ 1 CCA (11) A worm tooth surface is drawn based on Eq. (11) in Fig. 6. The ranges of parameters are q 2 rf 1; ra1\u00bd and h 2 p;p\u00bd . 3.3 Transformation Matrix of Coordinates. The transformation matrix is needed to transform the tooth surface of the gear to the coordinate system of the worm. Figure 1 shows the positional relationship of the teeth-skipped worm and the gear to be measured. The axes of the worm and the gear are straight lines in different planes. During the measurement, the angle between Zw-axis and Zgaxis is the working shaft angle P w. It is expressed as X w \u00bc b16b2 \u00fe eR (12) where b1 and b2 are the helix angles of worm and gear and eR is the misalignment of the shaft angle. The distance between Zw-axis and Zg-axis is the working center distance aw. It is determined as aw \u00bc rw1 \u00fe rw2 \u00bc r1 \u00fe r2 \u00fe ea (13) where r1 and r2 are the radiuses of reference circles of worm and gear and ea is the misalignment of the center distance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure7.9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure7.9-1.png", + "caption": "Fig. 7.9 Capacities within a slip ring asynchronous machine", + "texts": [ + " The shaft voltage over the bearing ub is related to the common mode voltage uCM by the following relationship: brfwr wr CM b 2 CCC C u u \u22c5++ = . (7.7) Whereas the voltage is properly defined by equation 7.7, the current is not. Indeed the current is a result of different phenomena [7.1]. In case of a slip ring supply for the rotor winding, a capacitive shaft voltage dominates any other kind of shaft voltage in general, if the shaft is not properly grounded. This case is evident for slip ring turbo-generators as well as for slip ring asynchronous machines. Main capacities are given in fig. 7.9. The strong capacitive coupling is on such a level, that any machine with direct converter supply to the rotor has to have proper rotor grounding. The voltage can be calculated according equation 7.8: brfrw2 rw2 CM2 b 2 CCC C u u \u22c5++ = . (7.8) 7.3 Measurements of Current Path and Voltage Transients 119 Shaft voltages of several hundred volts occur during the operation of turbogenerators or other large electrical machines with static frequency converters, if certain conditions are fulfilled. These conditions can be elaborated by measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000826_s1052618814050136-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000826_s1052618814050136-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " The \u201cDELTA\u201d manipulator is one of the most famous schemes of such type [7]. Though all the motors of this manipulating mechanism are rotational, the out put element does not change its orientation. Because the intermediate elements of mechanisms have a light load, this manipulator has a quite high operating speed. A number of similar devices were developed under the guidance of Vincenzo Parenti Castelli [8]. The well known manipulating mechanism \u201cORTHOGLIDE\u201d [9, 10] contains three kinematic chains and three translational motors (Fig. 1). Moreover, every chain has the parallel crank mechanism which provides stability of the output element orientation. This scheme is in many ways similar to the \u201cDELTA\u201d manipulator, but the rotational motors were changed for translational ones. A considerable number of publications are devoted to spherical manipulators where the output ele ment performs the rotational movement only [11, 12]. This is important for orientation of antennas, tele scopes and different types of tools. The actual task is making the kinematic interchange that divides the generalized coordinates for han dling the translational motion and handling the rotational motion", + " 5 2014 SYNTHESIS OF MECHANISMS OF PARALLEL STRUCTURE 379 Therefore, the task for the development of the structural schemes for the manipulating mechanisms of the parallel structure was set up when the rotation would transfer via paral lel crank mechanisms which are a part of each kinematic chain. Let us consider the basis for the structural synthesis of mechanisms being the evolution of the \u201cORTHOGLIDE\u201d robot with due account of additional degrees of freedom. Each kinematic chain of the mechanism of this robot contains the flat parallel crank mechanisms (Fig. 1). Besides, they have two turning kinematic pairs with axes perpendicular to axes of the parallelogram i.e., each kinematic chain contains at least two universal (two degrees of freedom) joints. This fact may be used for transmitting rotation. It is known [16] that the kinematic chain, including two degrees of free dom joints with parallel axes, may be equipped with one or two additional chains. Given that, the kinematic chain is able to transfer rotation with a transmission ratio that equals one and keeps the mutual orientation of the input and output chains" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000843_techsym.2014.6808086-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000843_techsym.2014.6808086-Figure1-1.png", + "caption": "Fig. 1 Per-phase equivalent circuit of a three-phase SEIG", + "texts": [ + " Identifying explicitly the machine parameters that affect the values of the excitation capacitor and generated frequency. 2. Addressing different conditions no-load, resistive and inductive loads. To confirm the validity of the proposed analytical approach, values of obtained capacitance are taken for both simulation and experimental. Voltage build up process is described in this paper in d-q stationary frame. To determine minimum excitation capacitance the per-phase equivalent circuit of a three phase SEIG with an R-L load and an excitation capacitor is used as shown in Fig. 1. Rs, X1s, Rr, Xlr, Xm, Xc, RL and XL are per phase stator resistance, stator stator leakage reactance, rotor resistance, rotor leakage reactance, magnetizing reactance, excitation capacitive reactance, load resistance and reactance, respectively. The reactances in the equivalent circuit are calculated at base frequency. F and \u03c5 are p.u frequency and speed. Is, Ir, Im, Ic, and IL are per phase stator, rotor, magnetizing, capacitor and load currents respectively. R TS14P02 370 978-1-4799-2608-4/14/$31.00 \u00a92014 IEEE 408 From Fig. 1 and under steady state: 0s tI Z (1) where ( // ) ( // )t s L c L rZ Z Z Z Z Z / ,s s sZ R F jX /L L LZ R F jX 2/c cZ jX F , m mZ jX /( )r r rZ R F jX Since for self-excitation 0,Is thus Zt in (1) is equal to zero. Zt is given by (2), ( )( ) ( ) ( )t m r L C L C m r m r L CSZ Z Z Z Z Z Z Z Z Z Z Z Z Z (2) A. No- load For no-load ,ZL after substituting all the values of impedances given in Fig. 1 frequency of generated voltage and minimum value of excitation capacitance of SEIG can be obtained from (3). ( ) ( ) 0c r m s r m r mZ Z Z Z Z Z Z Z (3) Separating real and imaginary part of (3) and equating to zero. The minimum value of capacitance required for self excitation of SEIG and frequency of the generated voltage of SEIG at no- load is obtained from imaginary and real part of (3) and presented in (4) and (5). 2 max ( )( )s r c ls m r R R F F X X X F F R (4) 2 3 2 2(2 ) (2 ) ( ( ) ) ( ) 0 ls m ls mls ls c m ls s r c m ls X X X F X X X F X X X R R F X X X (5) In no-load operation of SEIG, the machine slip is almost zero, and the pu frequency of the generated voltage nearly equal to the pu rotor speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003350_tmag.2020.3010617-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003350_tmag.2020.3010617-Figure4-1.png", + "caption": "Fig. 4. Conceptual side view for calculating the stacking factor.", + "texts": [ + " 0018-9464 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. CU-01 4 III. ELECTROMAGNETIC PERFORMANCES CONSIDERING RADIALLY LAMINATED CORE A. Stacking Factor with Radial Lamination The outer and inner cores have a radial lamination structure to secure manufacturing possibilities and minimize the effects of eddy currents generated in electric steel. However, as shown in Fig. 4, the radial lamination structure reduces the effective flux path and consequently reduces the electromagnetic performances. Therefore, the calculation of electromagnetic performances proposed in this paper considers the reduced area of the core by defining the stacking factor and applying it to the performance calculation. The stacking factor is the ratio of the effective area resulting from stacking to the total area and can be expressed as follows [14]: , , core effective sf core total A k A (15) where Acore,effective and Acore,total are effective area and total area of the core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000302_978-1-4471-4141-9_29-Figure29.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000302_978-1-4471-4141-9_29-Figure29.4-1.png", + "caption": "Fig. 29.4 SwarmItFIX developed end effectors", + "texts": [ + " Two head designs were developed in SwarmItFIX, magneto-rheological fluid head (MRF) and vacuumclustered incompressible grains head (sand) [3, 19]. The main common requirements are: \u2022 adhesion at one side of the workpiece only (no geometric or closure grasp possible) \u2022 non magnetic adhesion possible (non-ferromagnetic workpiece) \u2022 Glues and sticky tapes not feasible (smooth workpiece surface with delicate protective layers not to be damaged) \u2022 high adhesion force per unit area [11]. The MRF head, Fig. 29.4a and c, has a triangular shape to minimize the distance between its perimeter and the tool. A crown of small pistons communicating through a network of channels are distributed along the perimeter of the head. The pistons are semi drowned in a room full of magneto-rheological fluid and are constrained to shift along the head axis. When the magneto-rheological fluid is not magnetized its viscosity is low enough to let the pistons translate copying the workpiece shape along the head perimeter, once the shape is copied, a pneumatic cylinder drives a permanent magnet close to the network of channels containing the MR fluid increasing its viscosity nearly to solid state blocking all pistons from further translation. Vacuum is applied inside the perimeter of the head reproducing the function of a suction cup. The orientation of the head triangle is commanded by a dedicated degree of freedom. The sand head, Fig. 29.4b and d, is of circular shape not needing an additional rotation to orient it with respect to the workpiece. The adaptation is achieved putting granular material in an compliant envelope and the adhesion used is vacuum. The head is covered by an outer lip that gives the necessary sealing for the vacuum to be produced. Initially the sand inside the head\u2019s body can take any shape just by pushing it against the workpiece; the vacuum is then generated compacting the grains in the envelope and generating head-workpiece adhesion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002017_1.4033364-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002017_1.4033364-Figure17-1.png", + "caption": "Fig. 17 Spring models in ANSYS", + "texts": [ + " In order to simulate this effect, stiffness is imported as tabular data as shown in Fig. 16. When a negative deflection occurs in the spring, it applies a compressive force. However, when the spring length increases\u2014meaning positive deflection occurs, no force is applied at the contact. In order to make the solver converge to a solution, negligibly small value of force is used (i.e., 100 N for 1 mm deflection). In total, 290 springs are connected equidistantly between inner and outer raceways (Fig. 17). Then, axial and tilting moment loadings are applied to the bearing to obtain roller contact forces. Analyses are performed according to the test procedure given in Sec. 3. Two different types of analysis are performed: pure axial loading analysis combined loading analysis 4.3 Load Distribution Analysis Results. In pure axial loading analysis, it has been observed that the load is distributed among the rollers located in the upper row. No force is exerted on the rollers located in the lower row, since all the load is supported by the rollers in the upper row" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure1-1.png", + "caption": "Fig. 1 The sketch of orthogonal turn-milling", + "texts": [ + " Present researches on turnmilling machining mainly focus on the cutting path and the surface roughness of workpiece [2\u20136]. With the increasing demands of manufacturing, cost spent on cutters during turnmilling machining is becoming larger and larger. So cutter wear and cutter life must be considered. The turn-milling operations can be divided into two groups by Schulz et al. [1]: orthogonal and co-axial. The focus of this paper is orthogonal turn-milling. The movement of orthogonal turn-milling includes a rotation motion of the workpiece, a feed motion, and rotation motion of the cutter as shown in Fig. 1. The cutter used in turn-milling is a cylinder with several blades equally distributed along the circumferential direction around the central axis. Every cutter blade shares the same cutting condition and wear situation. If the cutter blade is worn badly during machining, the surface of workpiece will also be destroyed. In order to avoid this case occurring, conservative method is adopted usually, that is, replace the cutter blade before badly worn according to experience. In practice, we found the worn part of cutter blade is concentrated on a small section around the corner, while the other sections of the cutter blade are intact and still have potential to use" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000807_s12541-015-0062-9-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000807_s12541-015-0062-9-Figure5-1.png", + "caption": "Fig. 5 Fixture (a) Diagram of the fixture (b) Photos of real products", + "texts": [ + " The elastic deformation is quite small, and the value is in the range of 0.1 to 20 microns. This brings high performance requirement for the sensor, the measurement system and the fixture. In this paper, the sensor which is specially designed to test the deformation of the bearing outer ring REBAM\u00ae (Rolling Element Bearing Activity Monitor) is chosen. The sensor can accurately measure the deformation of the bearing outer ring and transfer it into current signals. The REBAM\u00ae sensor can work correctly only a certain space around is retained. As shown in Fig. 5(a), the diameter of the open hole, denoted as D, is three times large of the diameter of the probe. The full fixture is illustrated bu the photos in Fig. 5(b), which is specially designed for the evaluation of the quality of the roller bearing. The advantage of the fixture is that it is easy to load and unload the testing bearings, which is quite useful for the batch evaluation of the bearing quality in the production line. To test and verify the feasibility of the proposed sensor and the experimental device, five sets of bearings with different quality problems are selected as test samples. There are four qualified bearings, four bearings with abnormal roughness on the outer ring, four bearings with abnormal roughness on the inner ring, four bearings with bruise on the outer ring, and four bearings with bruise on the inner ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002816_1350650119897471-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002816_1350650119897471-Figure7-1.png", + "caption": "Figure 7. MDOF gear dynamic model.", + "texts": [ + " MDOF model of gear pair During gear meshing, the friction force acts on the OLOA direction. To reflect the friction behavior of gears accurately during meshing and study the coupling relationship between the friction characteristics of the tooth surface and the dynamic characteristics of the gear system, the influence of OLOA direction should be added when considering the translational degree of freedom of gears. Therefore, the MDOF gear dynamic model is adopted to calculate the dynamic response of the gear system. Figure 7 shows the schematic of the dynamic model. On the basis of the diagram, the governing equations are listed as follows I1 \u20ac 1 \u00fe rb1 km\u00f0rb1 1 rb2 2 \u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2 \u00fe _y1 _y2 _e\u00de \u00bc T1 \u00fe XN i\u00bc1 Lpi i km\u00f0rb1 1 rb2 2 \u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2 \u00fe _y1 _y2 _e\u00de \u00f05\u00de m1 \u20acy1 \u00fe cy1 _y1 \u00fe ky1y1 \u00bc km\u00f0rb1 1 rb2 2 \u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2 \u00fe _y1 _y2 _e\u00de \u00f06\u00de m1 \u20acx1 \u00fe cx1 _x1 \u00fe kx1x1 \u00bc XN i\u00bc1 i km\u00f0rb1 1 rb2 2 \u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2 \u00fe _y1 _y2 _e\u00de \u00f07\u00de I2 \u20ac 2 rb2 km\u00f0rb1 1 rb2 2\u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2\u00fe _y1 _y2 _e\u00de \u00bc T2 XN i\u00bc1 Lgi i km\u00f0rb1 1 rb2 2\u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2\u00fe _y1 _y2 _e\u00de \u00f08\u00de m2 \u20acy2 \u00fe cy2 _y2 \u00fe ky2y2 \u00bc km\u00f0rb1 1 rb2 2 \u00fe y1 y2 e\u00de \u00fe cm\u00f0rb1 _ 1 rb2 _ 2 \u00fe _y1 _y2 _e\u00de \u00f09\u00de m2 \u20acx2 \u00fe cx2 _x2 \u00fe kx2x2 \u00bc XN i\u00bc1 i km\u00f0rb1 1 rb2 2 \u00fe y1 y2 e\u00de\u00f0 \u00fecm\u00f0rb1 _ 1 rb2 _ 2 \u00fe _y1 _y2 _e\u00de \u00f010\u00de where N is the number of tooth pairs meshing at the time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure11-1.png", + "caption": "Fig. 11. External actions on the Waveboard.", + "texts": [ + " To obtain the equations of motion of the system, one resorts firstly to the well known Newton\u2013Euler equations, which can be written in matrix form as M ( q ) \u0308q + D T ( q ) \u03bb = Q v ( q , \u02d9 q ) + Q g ( q ) + Q M ( q ) + A \u03c4 ( q ) \u03c4, (37) where M ( q ) is the mass matrix, \u03bb is the Lagrange multipliers set, Q v is the generalized quadratic velocity forces vector, Q g is the generalized forces vector associated with the weight of the bodies of the Waveboard, Q M is the generalized forces vector due to the elastic torques generated by the torsion bar (hereafter denoted by M 2 el and M 5 el ) and the term A \u03c4 ( q ) \u03c4 corresponds with the generalized external forces vector due to the actuations, \u03c4 . Five actuations are going to be considered, as many as the number of degrees of freedom of the system: \u03c4 = ( F 2 x F 2 y M 2 x F 5 y M 5 x )T . (38) Actuations in Eq. (38) are shown in Fig. 11 . Note that forces F 2 y and F 5 y correspond with the lateral forces F R and F F described in Section 2 and depicted in Fig. 4 . Moreover, two external torques M 2 x and M 5 x are applied along the longitudinal axis of the Waveboard. Two important aspects should be highlighted. In the first place, the presence of the torsion bar leads to the appearance of the two elastic torques M 2 el and M 5 el , which are proportionate to the relative angle among the platforms, \u03b852 . They have the same modulus and opposite senses: M 2 el = \u2212M 5 el = k t \u03b852 , (39) with k t being the torsional stiffness of the torsion bar. It is important to emphasize that M 2 el and M 5 el must not be confused with the external torques M 2 x and M 5 x previously defined. Finally, the presence of the forward force F 2 x (drawn with dashed line in Fig. 11 ) must be clarified. In the explanation of the Waveboard maneuvering of Section 2 , it was not mentioned the existence of any forward force that helped the rider to move forward. Nevertheless, it has been included in the model in order to ensure that the system is not underactuated, taking into account that its value in realistic trajectories should be close to zero, or at least much lower than the lateral forces. On the other hand, the dynamic equilibrium equations must be enlarged with the constraints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.2-1.png", + "caption": "Figure 10.2 A 3PRR+ 3RPR planar parallel manipulator with an intermediate platform.", + "texts": [ + " In that case, the third kinematic loop KL-3 = B1B3A3A1B1 becomes dependent on IKL-1 and IKL-2 because it is equal to their union, i.e. KL-3 = (IKL-1)\u222a (IKL-2). Alternatively, the two independent loops could be taken as IKL\u2032-1 = B1B2A2A1B1 and IKL\u2032-2 = B1B3A3A1B1. In that case, the third loop KL\u2032-3 = B2B3A3A2B2 would again be dependent on IKL\u2032-1 and IKL\u2032-2 because these three loops would still be related to each other so that IKL\u2032-2 = (IKL\u2032-1)\u222a (KL\u2032 \u2212 3). Example 10.2 3PRR + 3RPR Planar Parallel Manipulator The parallel manipulator shown in Figure 10.2 consists of three platforms. The fixed platform is the base link 0 that accommodates the guideways for the three sliders, i.e. the links 1, 2, and 3, which function as the feet of the manipulator. Their pivot centers are D1, D2, and D3. The terminal platform (link 14) carries the end-effector (QP). The intermediate platform (link 7) accommodates six bearings. The centers of the upper bearings are B1, B2, and B3. They belong to the legs B1A1, B2A2, and B3A3 that consist of the link pairs {8,11}, {9,12}, and {10,13}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003266_1350650120936856-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003266_1350650120936856-Figure1-1.png", + "caption": "Figure 1. (a) Schematic of a six-pocket hybrid journal bearing system and (b) different form of geometrically irregular JBS.11", + "texts": [ + " The use of CSL results in an improvement in bearing performance. Hence, the degradation in the bearing performance due to journal irregularities may be recovered up to an extent by using CSL instead of Newtonian lubricant. Further, none of the study in literature illustrates the influence of CSL on the sixpocket hybrid irregular JBS. In view of this, a study is planned to examine the effects of CSL on the characteristics parameters six-pocket hybrid irregular JBS with capillary restrictor as depicted in Figure 1(a). The outcomes of the study may be useful for an efficient and durable design of JBS. The flow behavior of CSL in a finite width JBS is modeled by using Stokes theory of CS fluid. The flow of a CS fluid is governed by the continuity and momentum equations given by Stokes as follows14,16 r:v \u00bc 0 \u00f01\u00de Dv Dt \u00bc rp\u00fe B\u00fe 1 2 r C\u00fe r2v r4v \u00f02\u00de To govern the CSL flow in bearing system as per Cartesian coordinates indicated in Figure 1, the expansion of equations (1) and (2) yields as follows16,24 @u @x \u00fe @v @y \u00fe @w @z \u00bc 0 \u00f03\u00de @p @x \u00bc @2u @z2 @4u @z4 \u00f04\u00de @p @y \u00bc @2v @z2 @4v @z4 \u00f05\u00de @p @z \u00bc 0 \u00f06\u00de Incorporating appropriate boundary conditions, equations (4) and (5) are solved to determine the velocity components (u and v)16,19 as follows u \u00bc Uz h \u00fe 1 2 @p @x z z h\u00f0 \u00de \u00fe 2l2cs 1 cosh 2z h 2lcs cosh h 2lcs 8< : 9= ; 2 4 3 5 \u00f07a\u00de and v \u00bc 1 2 @p @y z z h\u00f0 \u00de \u00fe 2l2cs 1 cosh 2z h 2lcs cosh h 2lcs 8< : 9= ; 2 4 3 5 \u00f07b\u00de where lcs \u00bc p The expressions of velocity components u and v are used in equation (3) and then integrated with respect to z", + " Further, incorporating the usual boundary conditions, the modified Reynold\u2019s equation yields as follows19 @ @x h3 12 \u2019cs h, lcs\u00f0 \u00de @p @x \u00fe @ @y h3 12 \u2019cs h, lcs\u00f0 \u00de @p @y \u00bc U 2 @h @x \u00fe @h @t \u00f08\u00de where \u2019cs h, lcs\u00f0 \u00de \u00bc 1 12l2cs h2 \u00fe 24l3cs h3 tanh h 2lcs \u00bcCS function. The non-dimensionalization using suitable parameters yields the governing Reynolds equation as follows22 @ @ h 3 12 \u2019cs h, lcs @ p @ ! \u00fe @ @ h 3 12 \u2019cs h, lcs @ p @ ! \u00bc 2 @ h @ \u00fe @ h @ t \u00f09\u00de where \u2019cs h, lcs \u00bc 1 12 l2cs h 2 \u00fe 24 l3cs h 3 tanh h 2 lcs ( lcs ! 0 means that lubricant will be Newtonian lubricant). The oil-film profile ( h) in a finite width geometrically irregular JBS shown in Figure 1(b) yields as follows6,10,11 h \u00bc 1 XJ cos ZJ sin \u00fe hgi \u00f010\u00de where hgi \u00bc a1 sin 2l barrel journal, a2 sin 2l bell-mouth journal, 0:5 a3 sin n \u00fe \u00f0 \u00de undulated journal 8>< >: 9>= >; \u00f010a\u00de Finite element method formulation Finite element method (FEM) is applied to compute the oil-film pressure ( p) in the governing equation (9). The oil-film domain is discretized into subdomains by means of fournoded quadrilateral iso-parametric elements. The oil-film pressure is considered to be linearly distributed over an element and yields as follows2,22,23 p \u00bc Xnel j\u00bc1 pjNj \u00f011\u00de where Nj \u00bc Lagrangian shape function and nel \u00bc no" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003336_cae.22305-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003336_cae.22305-Figure2-1.png", + "caption": "FIGURE 2 (a) Comparison of cutter/workpiece axial engagement in ball\u2010 and barrel\u2010end mills. (b) Geometry for the barrel\u2010end mill", + "texts": [ + " This section presents the basis of the force and topography models for barrel\u2010end mills and thus the underlying mathematics behind the training software. Oval or barrel\u2010shaped end mills are high\u2010performance tools for medium and finish milling operations. Most common diameters are often made of 3\u20134 flutes with a variety of coatings to adapt to different materials. The most significant feature of a barrel cutter is a reference circle that enables a wider cutter\u2010workpiece axial engagement than ball\u2010end mills (Figure 2). The radius of the circle, with values between 150 and 190mm, allows a more progressive tool entrance and reduces tool axial passes improving surface quality. For barrel tools, the main tool parameters are the diameter D2 (R2) of the envelope curve of the cutting edges, the reference diameter D that normally coincides with the shank diameter and the helix angle i0. See Appendix A for the definition of the main geometrical parameters. The model simulates the motion of the cutting\u2010edge points during the milling process", + " Bridging the gap between student instruction and advanced research: Educational software tool for manufacturing learning. Comput Appl Eng Educ. 2020;1\u201313. https://doi.org/10.1002/cae.22305 Cylindrical tools are sculptured with a constant helix angle (i0). However, tools with no cylindrical profile such as ball, nose or barrel\u2010end mills project this angle into a z\u2010dependent i(z) angle. On the other hand, the shift angle \u03c8(z) is related to the height of the edge element. For the discretization of the cutting\u2010edge points (white curves in Figure 2a), the following governing equation is obtained: r \u03d5 j z xi yj zk\u00af ( , , ) = ( \u00af + \u00af) + \u00af, (A1) where x= r(z)\u00b7sin\u03d5j, y= r(z)\u00b7cos\u03d5j and z= z. If the reference z= 0 is set at r(z)= 0.5D, the tool radius r(z) can be defined as: r z D z D D( ) = (0.5 2) \u2212 \u2212 0.5 ( 2 \u2212 ),2 2 \u22c5 (A2) where D is the tool diameter and D2 = 2R2, with R2 being the radius of the circle\u2010segment. Similarly, the approach angle varies with z and can be written as: \u03ba z D z D( ) = sin (2 (0.5 2) \u2212 / 2).\u22121 2 2 (A3) On the one hand, the program estimates the milling force components for a number of periods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000475_20140313-3-in-3024.00209-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000475_20140313-3-in-3024.00209-Figure2-1.png", + "caption": "Fig. 2.Photograph of our quad-rotor machine", + "texts": [ + " In order to derive a complete dynamic model, the rigid body dynamics and the effects of aerodynamics are studied in this section based on the reference paperDong et al (2013), which is modified in our work to account for variations in air-density variations with change in the altitude. Rigid body dynamics of the quad rotor UAV governs the response of attitude control. The expressions are derived in two coordinate systems: an inertial coordinates and a body fixed coordinates. The body fixed coordinates is defined as follows. As indicated in Fig. 2, the lever marked with black strip is chosen as the x axis, and the perpendicular lever is the x axis. Then the z axis is defined by the right hand rule. For inertial coordinates, the point where the quad rotor starts its flight is set as the origin, and an east-north-up orthogonal coordinate system is established by the right hand rule. With attitude angles defined as in Fig. 3, the transformation matrix R from inertial coordinates to body fixed coordinates as described in Dong et al(2013)is given by equation (1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002290_speedam.2016.7525820-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002290_speedam.2016.7525820-Figure2-1.png", + "caption": "Fig. 2. Structure of model A.", + "texts": [ + " The torque characteristics of the basic and proposed models are analyzed using the three-dimensional finite element method. We demonstrate that one of our proposed models has the same torque characteristics as that of the basic model. II. BASIC MODEL Fig. 1 shows the structure of the basic model with only NdFe-B magnets. This model has eight poles and 36 slots. The total volume of the Nd-Fe-B magnets is 58000 mm3. Tables I and II show the specifications of the basic model and Nd-Fe-B magnet, respectively. 978-1-5090-2067-6/16/$31.00 \u00a92016 IEEE III. PROPOSED MODELS Fig. 2 shows a new rotor structure, and the red arrows show the magnetized direction. Model A comprises two claw-pole-type iron cores which are called the upper and lower cores, a shaft of non-magnetic material, Nd-Fe-B magnets, a cylindrical ferrite magnet which is magnetized in axial directions, and auxiliary ferrite magnets. Table III shows the specifications of the ferrite magnet. This model is of the consequent pole type, allowing the total volume of the Nd-Fe-B magnets to be reduced by 57.5 % compared with that of the basic model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002272_2016-01-1958-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002272_2016-01-1958-Figure1-1.png", + "caption": "Figure 1. Automotive wheel bearing.", + "texts": [ + "ithin the automotive industry, there have been increased demands for compact, lightweight and maintenance-free wheel bearings. This has driven progression within wheel bearing research and development, specifically the Generation 3 bearing. Generation 3 bearings incorporate flanges into the bearing inner and outer rings for mounting directly to the vehicle knuckle and wheel. Figure 1 displays a Generation 3 wheel bearing and the mating and surrounding components. The flanges are subject to deformation as a consequence of fastening to the adjoining components. Deformation must be considered during design and testing. [1], [2] The outer ring of the Generation 3 wheel bearing is fastened to the vehicle knuckle through a bolt-induced clamping force which causes distortion of the outer ring, specifically, at the seal mounting location and raceways. If, by design and in the manufacturing process, the outer ring flange forms a concave shape, it will influence the distortion generated during the tightening process between the outer ring and knuckle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002397_icma.2016.7558663-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002397_icma.2016.7558663-Figure2-1.png", + "caption": "Fig. 2 North-East and body-fixed coordinate system", + "texts": [ + " This paper is intended to take advantage of GPI observer to estimate system status and general disturbance as well as to improve control accuracy, to ensure the FPSO be well positioned ,accurately and safety. Here, the planar three degrees of freedom (DOF) model in surge , sway and yaw is established to keep the FPSO\u2019s 789978-1-5090-2396-7/16/$31.00 \u00a92016 IEEE Proceedings of 2016 IEEE International Conference on Mechatronics and Automation August 7 - 10, Harbin, China heading and position. Generally, the ship model is divided into two parts: the kinematic model and hydrodynamic model [9]. Firstly, the North-East and body-fixed coordinate system are established in Fig 2. Planar motion mathematical model of FPSO vessel is shown below: ( ) ( ) ( ) ( ) v J M C v v D v v g w \u03b7 \u03c8 \u03c5 \u03b7 \u03c4 = + + + = + (1) where, ( )J \u03c8 is coordinate transformation matrix; [ ]= Tn e\u03b7 \u03c8 is the ship position state vector of the North- East coordinate system; [ ]= Tv u v r is the velocity vector of the body-fixed coordinate system; \u03c4 is general control forces; w is general environmental forces. M denotes ship additional mass inertia matrix; ( )C v denotes the ship Coriolis centripetal force matrix; ( )D v is hydrodynamic damping matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure4-1.png", + "caption": "Fig. 4. Maneuvering of the Waveboard.", + "texts": [ + " Subsequently, the rider turns both feet out and moves the torso in the opposite direction, describing the motion denoted by (2). Repetition of this sequence over time leads to the forward propelling of the Snakeboard. In the case of the Waveboard, the forward propelling results from the oscillatory motion of both decks through lateral actions exerted by the rider with their feet. In this way, the lateral forces press the platforms down, making them roll about the axis of the Waveboard, and each pasive caster wheel is pushed out to the opposite direction. Fig. 4 illustrates the previous ideas, with both platforms (rear board is denoted by R and forward by F ) twisted due to the application of rear and forward lateral forces F R and F F , respectively. Note that these actions generate a force couple, denoted by M , which is congruent with the steering directions of rear and forward wheels. A physical explanation for the forward motion can be given projecting lateral forces F R and F F on the steering directions of the wheels. In Fig. 4 one can see that these forces do work, being its projections F W R and F W F nonzero, contributing to the forward propelling. Thus, rolling both decks alternately (cyclic variation of angles \u03c6r and \u03c6f ) leads to the meandering trajectory shown in sequence (1)\u2013(4) of Fig. 5 . It is important to emphasize that rear and forward forces are not applied simultaneously, existing a gap between them. In configuration (1) of Fig. 5 , the rider exerts a lateral rear force, twisting the rear platform, and the forward force is small", + " To obtain the equations of motion of the system, one resorts firstly to the well known Newton\u2013Euler equations, which can be written in matrix form as M ( q ) \u0308q + D T ( q ) \u03bb = Q v ( q , \u02d9 q ) + Q g ( q ) + Q M ( q ) + A \u03c4 ( q ) \u03c4, (37) where M ( q ) is the mass matrix, \u03bb is the Lagrange multipliers set, Q v is the generalized quadratic velocity forces vector, Q g is the generalized forces vector associated with the weight of the bodies of the Waveboard, Q M is the generalized forces vector due to the elastic torques generated by the torsion bar (hereafter denoted by M 2 el and M 5 el ) and the term A \u03c4 ( q ) \u03c4 corresponds with the generalized external forces vector due to the actuations, \u03c4 . Five actuations are going to be considered, as many as the number of degrees of freedom of the system: \u03c4 = ( F 2 x F 2 y M 2 x F 5 y M 5 x )T . (38) Actuations in Eq. (38) are shown in Fig. 11 . Note that forces F 2 y and F 5 y correspond with the lateral forces F R and F F described in Section 2 and depicted in Fig. 4 . Moreover, two external torques M 2 x and M 5 x are applied along the longitudinal axis of the Waveboard. Two important aspects should be highlighted. In the first place, the presence of the torsion bar leads to the appearance of the two elastic torques M 2 el and M 5 el , which are proportionate to the relative angle among the platforms, \u03b852 . They have the same modulus and opposite senses: M 2 el = \u2212M 5 el = k t \u03b852 , (39) with k t being the torsional stiffness of the torsion bar. It is important to emphasize that M 2 el and M 5 el must not be confused with the external torques M 2 x and M 5 x previously defined", + " Nevertheless, it is important to study the influence that these two basic design parameters have on the dynamics of the Waveboard. For that purpose, simulations varying \u03b8 c from 26 \u25e6 to 34 \u25e6 and k t from 100 to 400 N m rad for the winding trajectory in Eq. (56) are going to be performed. In the first place, the influence of \u03b8c is analyzed. Figs. 21 and 22 show the variation of the lateral forces F 2 y and F 5 y , being more important the effect of \u03b8 c in the case of F 2 y . It can be seen that these actions exerted by the rider are lower as the inclination angle increases. This fact can be easily explained resorting to Fig. 4 : as \u03b8 c rises, the work done by F 2 y and F 5 y increases, being their projections on the steering directions of the wheels larger. Nevertheless, in this analysis it is also important to point out that \u03b8 c cannot be unreasonably increased, since it would make the Waveboard completely unstable and impossible to master. Therefore, a compromise among comfortable maneuvering applying low lateral forces and stability is needed, being reasonable values of \u03b8 c around 30 \u25e6. With respect to the torques M 2 and M 5 , variation of \u03b8 c does not mean a substantial change, as it is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003589_1350650120962973-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003589_1350650120962973-Figure6-1.png", + "caption": "Figure 6. A10 Test gears having different face widths.38", + "texts": [ + " The dynamic normal load and the meshing gear power loss results are predicted through an iterative loop based on the convergence criterion. On the whole, the iterative loop mainly includes the illustrated steps in Figure 5. Extra details of some steps are given in authors previous work32). Several measurements were performed on the FZG test rig to validate the dynamic proposed models. More details regarding the experiments for the measurement of the spur gears power loss is described in authors previous work.32 All the following simulated and experimental results are obtained for FZG A10 spur gear pairs (see Figure 6). The geometric parameters of this type of gear are detailed in Table 2. The 75W90-A lubricant is a typical axle gear oil recognized as both rear axle and manual transmission fluid. It is selected since it represents a lower m value compared to the other axle lubricants suggesting that lubricant additives might play an important role on power loss of a gear pair as shown in previous experimental work.38 The selected lubricant whose properties listed in Table 3 is used under an operating temperature of 80 C in this study that considers different operating conditions listed in Table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000533_tmag.2015.2442245-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000533_tmag.2015.2442245-Figure4-1.png", + "caption": "Fig. 4. Experimental apparatus.", + "texts": [ + " In this configuration, the mass of the actuator was 21.8 g. This magnetic actuator was capable of reversible motion, rotary motion, and turning in the pipe. The principle of linear locomotion is considered first. The actuator was held in the pipe by initial supporting force F using compound material A, as shown in Fig. 3. SMA coils C and D were assumed to contract, and compound materials A and B were assumed to be open and completely closed, respectively. An experiment was conducted using the apparatus shown in Fig. 4. The actuator was driven using the same two signal generators of two channels and three amplifiers. When a sinusoidal electric current was applied to the electromagnet, the displacement of the vibration component became synchronized in the current waveform. During measurement, the displacements of vibration components 1 and 2 were synchronized by phase adjustment of the signal generator. All the vibration components were driven at the same frequency. The driven frequency of the actuator was 211 Hz. The vibration displacement of the moving actuator was measured using a laser displacement meter and a fast Fourier transform analyzer, and the supporting force P was measured using a force pickup and a charge amplifier. During measurement, the pipe was cut in half, and one half of the pipe was attached to the force pickup, as shown in Fig. 4. The initial supporting force F provided by compound material A and the weight of the actuator cancel each other, and only the change in the supporting force P by the vibration was measured. When the actuator was inserted in a pipe having an inner diameter of 35 mm, the angle \u03b8 was approximately 45\u00b0. Fig. 5 shows the response of the displacement for the vibration component and supporting force P acting on the pipe. When the vibration component was displaced by amplitude A in the +z-direction, supporting force P remained essentially unchanged" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003202_0142331220921024-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003202_0142331220921024-Figure9-1.png", + "caption": "Figure 9. Inverted pendulum system.", + "texts": [ + " Figure 8 shows the norm of adaptive parameter vectors uf , ug, ur and up and it is clear that all adaptive parameters are bounded. According to the obtained results, the proposed adaptive scheme guarantees the boundedness of all the signals of the closed loop system, which validate the theoretical results. The obtained results can be improved by using optimization techniques to find the optimal values of the design parameters. Example 2 In this example, the tracking control of the benchmark control problem of inverted pendulum, shown in Figure 9, is investigated. Using the proposed adaptive fuzzy controller, the obtained results are compared to those obtained by adaptive backstepping control approaches proposed in Zhou et al. (2004) and Zhou et al. (2007), respectively. The dynamic equations of the inverted pendulum system are given by (Wang, 1996) _x1 = x2 _x2 = gsinx1 mlx2 2 cosx1sinx1 mc +m l 4 3 m(cosx1) 2 mc +m +Df x\u00f0 \u00de 24 35+ cosx1 mc +m l 4 3 m(cosx1) 2 mc +m +Dg x\u00f0 \u00de 24 35w u\u00f0 \u00de+ d t\u00f0 \u00de y= x1 8>><>>: \u00f070\u00de where x1 = u and x2 = _u represent the angular position and the angular velocity, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003417_metroaerospace48742.2020.9160076-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003417_metroaerospace48742.2020.9160076-Figure2-1.png", + "caption": "Fig. 2. Exampe of sensorized support for a 3D printing head.", + "texts": [ + " This is an innovation of great interest especially in view of the status of the process. The suggested approach is an in-line testing. During the ALM production procedure, a series of built in test equipment should monitor the component during the construction, to gather data on the transient elements experienced by the part. Figure 1 shows a basic schema of this concept [33-40]. The basic idea introduced in this work is to monitor the process acquiring a timestamped measured from different sensors (Fig. 2). Correlating the accelerometric information and the G-CODE, to each timestamp can be joined a precise point of the handwork. In this way, in presence of a local defect of the printed object, it is possible to analyze the acquired data to understand how the printing process has been induced the defect. 529 Authorized licensed use limited to: University of New South Wales. Downloaded on September 20,2020 at 13:15:46 UTC from IEEE Xplore. Restrictions apply. Furthermore, by printing different replicas of an object, it is reasonable to assume that the different artefacts have similar characteristics if the sensors record similar values", + " Correlating the IMU acquisition and the estimated time evolution of the acceleration allows the time alignment of the \u201csensors domain\u201d and of the \u201cG-CODE domain\u201d. This alignment make possible the enrichment of the time based sensors data with the spatial position on the GCODE route. This information can be stored in a DB that allows to rapidly access to the sensors data acquired from a spatial model. The proposed workflow has been tested using the architecture showed in Fig. 6. The sensors group (showed in Fig. 2) contains the following sensors: \u2022 Atmel BNO055 (9-axes IMU); \u2022 2 Pi NoIR Raspberry Camera + a 40\u00b0 optics (video cameras); \u2022 FLIR Lepton 3.5 with radiometry (thermal camera). A dedicated Raspberry Pi 3 B+ board manages each sensor. In this way, for each sensor maximum performances are assured. A PC that host the DB communicates with the Raspberry board through an ethernet switch. In order to guarantee the same time base for the acquisition, a Real Time Clock (RTC) has been used. The RTC has been connected to one of the Raspberry board programmed as NTP server" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003445_j.procir.2020.05.186-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003445_j.procir.2020.05.186-Figure1-1.png", + "caption": "Fig. 1. test specimen with different geometrical features", + "texts": [ + " Therefore the CAD-model has to be deformed. 4. Validation and Discussion of the Results To validate the measuring method, a simplified specimen was constructed in a first step. This specimen includes some of the common geometric features. It contains different drill holes, a truncated cone, a circular ring, a chamfer, plains and a free form surface. The test specimen was milled on a five-axis milling machine (GROB G350, 1st Generation, GROB GmbH & Co. KG, Mindelheim, Germany) and it is shown in figure 1. Afterwards, the number of measuring points had to be defined. Except for the free form surface, values from literature according to [16] were used to characterize the different geometric features. The numbers of the applied points are listed in table 1. The calculation of the measuring point positions according to the Hammersley sampling strategy was done in MATLAB\u00ae (The MathWorks Corporation, Natick, USA). In Figure 2 the realization of the Hammersley sampling strategy of the contour for the free form surface is shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003635_s11249-020-01356-z-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003635_s11249-020-01356-z-Figure1-1.png", + "caption": "Fig. 1 Schematic of in\u00a0situ measurement using Raman spectroscopy", + "texts": [ + " Furthermore, the accuracy of temperature measurement on the sliding surface using Raman spectroscopy has not been investigated. Therefore, we conducted temperature measurements of the sliding surface using the three independent parameters of the Raman spectra. Friction tests were performed using a friction test rig comprising a stationary sapphire hemisphere and a rotating disk of steel in dry conditions. The Raman spectrum was in\u00a0situ detected for estimating temperature through the sapphire hemisphere during the tests. Tribology Letters (2020) 68:116 1 3 Page 3 of 12 116 Figure\u00a01 shows a schematic of the tribometer and Raman spectroscopy used in the current study. The tribometer is a pin-on-disk type tester, which creates a point contact between a stationary sapphire hemisphere of diameter 10\u00a0mm, and a rotating steel disk of diameter 60\u00a0mm and thickness 3\u00a0mm. A Raman spectroscopy equipment was installed above the tribometer to measure the temperature of the contact area. An infrared thermometer was used to compare the measured temperature with the temperature measured by Raman spectroscopy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000828_2013-01-0993-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000828_2013-01-0993-Figure1-1.png", + "caption": "Figure 1. Semi-active suspension system for half-car model", + "texts": [ + " In this paper, the influence of road roughness on the ride performance criteria using half vehicle model is evaluated. Also, the effect of road roughness on the total vehicular resistance is discussed. Furthermore, the percentage power dissipation of passive suspension elements and semi-active suspension system, with continuously variable orifice damper, relative to rolling resistance power losses are also evaluated. A basic half-car suspension model equipped with a semiactive damper considered in this study is shown in Figure 1. The vehicle body itself is assumed to be rigid, and has degrees of freedom in the vertical and pitch directions. The body mass and pitch inertia are represented by Mb and Ib, respectively. Mw, ks, Ct and kt denote the wheel mass, the spring stiffness, the tyre damping coefficient and the tyre spring stiffness, respectively. Suffix f and r denote front and rear, respectively. The equations of motion for this model are given by; (1) (2) (3) (4) (5) (6) (7) The semi-active damper is theoretically capable of tracking a force demand signal independently of instantaneous velocity across it", + " A semiactive independent typecontinuously variable damper allows changing the damping force by controlling the orifice area. The control law which optimizes semi-active system performance provides a control force of the form; (12) It is acceptable to consider the coefficient of rolling resistance as a linear function of car speed. The rolling resistance of tyre on hard surfaces is primarily caused by the hysteresis in tyre materials due to the deflection of the carcass while rolling. Vehicular rolling resistance force for half vehicle model (F), Figure 1 traversing an uneven road, can be written in the form; (13) and The vehicular rolling resistance coefficient, f equals (ft+fd +fs)/G. Equations describing tyre rolling resistance present traversing an absolutely smooth road (Ft), the additional tyre rolling resistance forces due to the tyre dynamic distortion (Fd) and tyre rolling resistance forces due to suspension motion excited by road input (Fs) are described by Soliman [6]. The rolling resistance coefficient, f is equals F/G. The following assumption has been considered; the vehicle itself is treated as a rigid body, the mass of wheel and tire is concentrated at the tire centre for front and rear wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003813_ffe.13393-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003813_ffe.13393-Figure5-1.png", + "caption": "FIGURE 5 Boundary condition of global model and submodel", + "texts": [ + " In order to reduce the calculation scale, the fatigue life evaluation is presented based on the representative \u03a9 domain. The choice of dimensions of representative \u03a9 domain can lead to the convergence results for the fatigue life of gears with oxide inclusion. For the finite element analysis of gear static contact, the intermediate gear shaft should keep stationary as a driven gear. According to the real service conditions, three displacement boundary constraints are applied on the intermediate gear shaft of the global model, as shown in Figure 5. In detail, the first constraint limits the rotation of the gear shaft at face A (the installation position of output wheel), the second constraint fixes the radial displacement of the gear shaft at face B (the mounting position of bears) and the last constraint restricts the axial displacement at face C (one end face of the gear shaft). A driving torque is applied on the cut surface (marked as blue colour) of the low-speed gear to implement the rotation of gears along the z axis. And the results of the cut surface displacement obtained from the global model can be introduced as the input load of the submodel as shown in Figure 5. For FE simulations, software ANSYS is utilized to obtain the stress field of gears with oxide inclusion. A 3D 20-node homogeneous structural solid element (i.e. Solid186) is used to mesh the gears including base metal and oxide inclusion. The number of nodes and elements are 994 778 and 721 757 for the global model and 923 736 and 641 241 for the submodel, respectively. The element types selected to model the contact between gears are Targe170 and Conta173. In summary, there are three steps to evaluate the fatigue life of gears with oxide inclusion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.37-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.37-1.png", + "caption": "Figure 9.37 Second kind of position singularity of the manipulator.", + "texts": [ + "589) also implies that this singularity occurs if the position of the end-effector is specified in such a way that the wrist point coincides with the shoulder point. That is, x1 = x3 = 0 (9.597) The above equations express the noticeable feature of this singularity that the upper and front arms of the manipulator have equal lengths and the front arm is folded completely over the upper arm. However, this singularity cannot occur exactly because of the physical shapes of the relevant links and joints. This singularity is illustrated (approximately) in Figure 9.37. Note that it is the same as that of a regular Puma manipulator. Therefore, the two singularities have the same consequences concerning the indefiniteness of \ud835\udf032 as discussed in Section 9.1. In this case, the manipulator can have only one kind of position singularity, which is the same as the second kind of position singularity explained above in Section 9.7.5.1. (a) Angular Velocity of the End-Effector with Respect to the Base Frame Let \ud835\udf14 = \ud835\udf146. Then, Eq. (9.563) leads to the following expression" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001482_1.4029187-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001482_1.4029187-Figure4-1.png", + "caption": "Fig. 4 A dead center position for a single flier eight-bar linkage [4,17]", + "texts": [ + " For the complex linkage staying at the dead center position, the three noninput joints, I23, I35, and I15 must be collinear. This can be seen in Fig. 3, in which the three noninput joints are collinear. Example 3. Dead center positions for a single flier eight-bar linkage. Journal of Mechanisms and Robotics NOVEMBER 2015, Vol. 7 / 044501-3 Downloaded From: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use This is an example for the dead center positions of a single flier eight-bar linkage (Fig. 4) presented by Pennock and Kamthe [4]. Pennock and Foster [17,18] proposed a geometric method to obtain the secondary instant centers of the eight-bar linkage. With the input given through link 2, the equivalent four-bar linkage can be I12I23I34I14, which comes from the four-bar loop that contains the fixed link and the input link and is type I four-bar equivalent linkage. In addition, another type of equivalent four-bar linkage can be I12I26I65I51, which comes from the four links 1, 2, 6, and 5 and consists of the corresponding four instant centers I12, I26, I65, and I51(I15). I12 is regarded as the input joint. With the method mentioned above, it is easy to determine the dead center positions of the eight-bar linkage. If any equivalent linkage is at dead center positions, the whole linkage is at dead center positions. With the input given through link 2 or I12, if the three passive joints I23, I34, and I41 of the type I equivalent four-bar linkage I12I23I34I41 are collinear, the linkage is at dead center position, as shown in Fig. 4. The three passive joints I26, I65, and I15 of the type II equivalent four-bar linkage I12I26I65I15 are automatically collinear. The same result as Ref. [4] can be obtained with the proposed method. Another case is that if one of the type II equivalent four-bar linkage is at dead center positions, the whole linkage is at dead center positions while the type I four-bar linkage is not necessarily collinear, as shown in Fig. 16 in Ref. [4]. Example 6 will explain this observation. Example 4. Dead center positions for a symmetric eight-bar linkage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003519_ccc50068.2020.9188831-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003519_ccc50068.2020.9188831-Figure1-1.png", + "caption": "Fig. 1: The coordinate system ofthe QSL system", + "texts": [ + " (c o s^ s in ^ co s^ -s in ^ s in ^ ) TT \u0302 f disy y ~ Ui ^m m cos

0 leads to a contradiction, we show that an alternative admissible sequence, {v(t)} can be obtained such that \u2225v(t) \u2212 (t)\u22252 < V\u2113 can be achieved and hence complete the proof. Define the following quantities for 0 \u2264 \u03bb \u2264 1: v(t, \u03bb) := (1 \u2212 \u03bb)\u03c1(t) + \u03bbr(t) (34) Vb(t, \u03bb) := \u2225v(t, \u03bb) \u2212 r(t)\u22252 (35) learly, v(t, \u03bb) is the convex combination of \u03c1(t) and r(t). Also, t each t , Vb(t, \u03bb) is a convex quadratic function of \u03bb, 0 \u2264 \u03bb \u2264 1 with the upper bound (1 \u2212 \u03bb)V (t). Choose \u03bb\u2113(t) such that (1 \u2212 \u2113(t))V (t) = V\u2113, see Fig. 3. Since V (t) \u2265 V\u2113 \u2265 0, it follows that \u2264 \u03bb\u2113(t) \u2264 1 and that \u03bb\u2113(t) \u2192 0 as t \u2192 \u221e as V (t) \u2192 V\u2113. Now efine a \u03bbb(t) such that \u03bbb(t) > 0 for all t \u2265 0 and \u03bbb(t) \u2192 0. Specifically, let \u03bbb(t) := \u03bb\u2113(t) + \u03b4(t) where {\u03b4(t)} is a sequence with \u03b4(t) > 0 for all t \u2265 0 and \u03b4(t) \u2192 0. That \u03b4(t) > 0 can be found follows from r(t) = St \u03c9\u03040 with \u03c9\u03040 \u2208 int(W\u221e). Choose the sequence of {v(t)} as v(t, \u03bbb(t)). Since Vb(t, \u03bb) is a convex quadratic function of \u03bb, in the limit as \u03b4(t) \u2192 0, it follows that \u03bbb(t) \u2192 \u03bb\u2113(t) and Vb(t, \u03bbb(t)) \u2192 V\u2113(t)+ dVb(t,\u03bb) d\u03bb (\u03bbb(t)\u2212 \u03bb(t))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure12-1.png", + "caption": "Figure 12. Experimental setup to measure the static frictional resistance of each gear type, including that with conical passive rollers", + "texts": [ + "2, meaning the width of one curved tooth at its outer rim is thinner than the maximum width at its center, with a total difference of 0.4 mm. In this research, we performed various experiments to determine the features and advantages of the two prototypes. Initially, we performed experiments to determine the feature of static and dynamic frictional resistance of the gear with conical passive rollers. To quantitatively determine the static frictional resistances of various gear types, we used the experimental setup shown in Fig. 12. Here, constant pressing force is applied by the force gauge on the right because some constant load is necessary to measure the differences in frictional resistance among different gear types. The pressing force was increased from 40 to 100 N. In all experiments to measure the static and dynamic frictional gear resistances in this paper, including this one to measure the static frictional resistance of a gear with conical passive rollers, the ball bearings used to support the central shaft of the spur gear or the gear with passive rollers for sliding were eliminated. The gear and shaft were fixed to the supporting block with a set screw for precise and stable measurement of the translational force required to achieve sliding motion while constant pressing force is applied from the right. We increased the weight on the Kevlar wire on the pulley on the left of Fig. 12 continuously to drive the pinion and rack gear for the sliding motion of the spur gear or gear with passive rollers, and measured the translational force when the fixed gear started to slide. We compared the static frictional resistances of a gear with crowning (crowning gear) [4], a spur gear, a Teflon-coated spur gear, and a gear with conical passive rollers. In this paper, we use the term \u201cRoller gear\u201d to mean a gear with conical or flat passive rollers in the graphs. The teeth width of the crowning gear, spur gear and Teflon-coated spur gear is 3.85 mm, which is the same as the largest diameter of the conical passive roller at its bottom. These gears\u2019 module is 2. Instead of the omnidirectional gear, we used the rack gear shown on the right of Fig. 12 because the manufactured initial gear prototype with conical passive rollers has only one line of passive rollers, and could not slide constantly between the intermittent teeth of the omnidirectional gear. However, the rack gear\u2019s module of the teeth and its material (SUS303) are the same as the omnidirectional gear to be driven by this gear with passive rollers, so its mechanical sliding motion feature is almost the same as that of the omnidirectional gear [7]. We performed the experiment ten times under each condition, and plotted the average value on the graph in Fig", + " The force gauge on the right showed that under every condition of pressing force, the t-test revealed significant differences between the static frictional resistance of the gear with conical passive rollers and other gear types. The advantage of this conical passive roller against static frictional resistance was thus confirmed in this experiment. In this research, the dynamic frictional resistances of the gear with conical passive rollers and other ordinary gears were also examined. The experimental setup used for this experiment is shown in Fig. 14. To achieve the constant sliding motion between each spur gear and the wider rack gear, we used the same normal rack and pinion gear mechanism shown in Fig. 12 and drove it with a geared AC motor, RSF-14B-100-F100-24B-C, manufactured by Harmonic Drive Systems Inc. Instead of the omnidirectional gear, we used a rack gear with wider teeth under the spur gear or the gear with passive rollers for the same reason as the static frictional resistance experiment described in the former chapter. We used the same spur and crowning gears as in the experiment to measure static frictional resistance. In this sliding experiment, the input torque from the geared AC motor was kept constant (0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000498_s40632-015-0014-7-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000498_s40632-015-0014-7-Figure1-1.png", + "caption": "Fig. 1. X \u2212 Y inverted pendulum system", + "texts": [ + " The paper is organized as follows: Firstly, the state space modeling equations of the X \u2212 Y inverted pendulum system have been presented. Secondly, the state equations in the presence of disturbance and friction have been obtained. Thereafter, the designing of adaptive gain scheduling PID controller is explained and finally the simulation results under different conditions are presented followed by the conclusion of the paper and references. An X \u2212 Y inverted pendulum mounted upon a base has been shown in Figure 1. The stabilization of the inverted pendulum in the upright position depends upon the horizontal displacement of the base, which in turn depends upon the applied forces F x and F y in the xy plane. As shown in Figure 1, x and y are the positions of the base from the reference in the xy plane. l is the distance from the base to the mass center of the pendulum. M and m are the mass of the base and the pendulum respectively. \u03b8 and \u03c6 are the angles made by the inverted pendulum with the vertical axis, when the base moves in x and y directions respectively. The parameters of the inverted pendulum system are given in Table 1 The state space model of the X \u2212 Y inverted pendulum system as proposed by Wang (2011) is given by: (1) (2) (3) (4) (5) (6) (7) (8) From Equation (1) to Equation (8), it can be concluded that the X \u2212 Y inverted pendulum is a two input four output system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001436_0954406214523581-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001436_0954406214523581-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of a symmetrical flexible rotor system supported on the self-acting gas bearings with threeaxial grooves.", + "texts": [ + " An algorithm is proposed to calculate the nonlinear dynamic behaviors of the rotor system supported on the self-acting gas-lubricated bearings with three-axial grooves based on the precise integration method30 because the common algorithms cannot be used to calculate the nonlinear dynamic behavior of the gas bearing-rotor system with the time delay properties and the time delays control. A symmetrical flexible rotor system supported on the self-acting gas bearings with three-axial grooves is shown in Figure 1. O1 and O3 are the geometric centers of the gas bearing, O2 is the geometric center of the disk. The dynamic equation of the self-acting gas bearing-symmetrical flexible rotor system can be written as m \u20acx\u00fe c _x\u00fe kx \u00bc f\u00fe w\u00fe q \u00f01\u00de at MICHIGAN STATE UNIV LIBRARIES on February 14, 2015pic.sagepub.comDownloaded from where m is the mass matrix, c is the damping matrix, k is the stiffness matrix, x is the displacement vector, f is the nonlinear gas film forces vector, w is the weight load vector, and q is the unbalanced force vector where m \u00bc m1 0 0 0 0 m1 0 0 0 0 m2 0 0 0 0 m2 2 6664 3 7775 c \u00bc d 0 d 0 0 d 0 d 2d 0 2d 0 0 2d 0 2d 2 6664 3 7775 k \u00bc k 0 k 0 0 k 0 k 2k 0 2k 0 0 2k 0 2k 2 6664 3 7775 x \u00bc \u00bd x1 y1 x2 y2 T f \u00bc \u00bd fx1 fy1 0 0 T w \u00bc \u00bd 0 m1g 0 m2g T q \u00bc m1ex1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002949_s12541-019-00277-9-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002949_s12541-019-00277-9-Figure4-1.png", + "caption": "Fig. 4 Back-side mesh stiffness calculation model", + "texts": [ + " For any piece, the total deformation \u03b4i of the tooth along the transverse plane under the mesh force on the action line can be calculated as: Each deformation can be calculated from the gear tooth parameters using Eqs.\u00a0(1)\u2013(4) [20]. Based on the definition of stiffness, the drive-side mesh stiffness for any piece and the total mesh stiffness during engagement can be obtained based on Eqs.\u00a0(6) and (7), respectively. where the quantities for the contacted pieces i can be calculated from the length of contact line l, as shown in Fig.\u00a03. in which, As shown in Fig.\u00a04, the back-side mesh stiffness can be calculated from the symmetry of the NTCIG gear teeth as: (4) F = arccos(rg\u2215rF) (5) i = Br12 + Bt12 + S12 + G12 + w12 + pV (6)ki = FNi i (7)kd = FN = \u2211 ki (8)l = \u23a7 \u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23a9 \ud835\udf03 \u2212 \ud835\udf03 1 \ud835\udf03 2 \u2212 \ud835\udf03 1 \u2217 b cos \ud835\udefd \ud835\udf03 1 < \ud835\udf03 \u2264 \ud835\udf03 2 b cos \ud835\udefd \ud835\udf03 2 < \ud835\udf03 \u2264 \ud835\udf03 3 b cos \ud835\udefd \u2212 \ud835\udf03 \u2212 \ud835\udf03 3 \ud835\udf03 4 \u2212 \ud835\udf03 3 \u2217 b cos \ud835\udefd \ud835\udf03 3 \u2264 \ud835\udf03 < \ud835\udf03 4 (9) 1 = tan F 1 \u2212 + 4x 1 tan 0 2z 1 \u2212 tan 0 \u2212 b tan 2r 1 2 = tan F 1 \u2212 + 4x 1 tan 0 2z 1 \u2212 tan 0 + b tan 2r 1 3 = tan k1 \u2212 + 4x 1 tan 0 2z 1 \u2212 tan 0 \u2212 b tan 2r 1 4 = tan k1 \u2212 + 4x 1 tan 0 2z 1 \u2212 tan 0 + b tan 2r 1 Based on the mesh stiffness calculation for the NTCIG derived above, the computation can be performed with the detailed gear parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure13-1.png", + "caption": "FIGURE 13. TEST FIXTURE FOR LEAKAGE AND PRESSURE CAPABILITY TESTING FOR VBD BRUSH SEAL.", + "texts": [ + " It is a high\u2013 pressure, high\u2013temperature rig with capability to adjust the upstream\u2013pressure as well as back\u2013pressure to simulate a variety of pressure differentials and pressure ratios. The design is similar to that described by Aksit et.al.[8], with the exception that the Coriolis flowmeter is placed in the exhaust line to get accurate readings for a high turn\u2013down ratio. Figure 12 shows the flow system layout of the leakage test rig. The fixture for mounting a segment of the seal into the rig is show in Figure 13. The rig and test conditions are described in brief in Table 2. The test was conducted in two stages. In the first stage, the downstream pressure, Plo was kept constant and the upstream pressure, Phi, was increased to 4.48 MPa (650 psig) by 0.69 MPa (100 psi) increments. In the second stage, Phi was raised to 4.48 MPa (650 psig) and Plo was decreased in steps of 0.69 MPa (100 psi), as shown in Figure 14. The leakage through the seal, , the upstream pressure, P1, downstream pressure, P2, and temperature, T1, are measured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002308_msf.862.11-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002308_msf.862.11-Figure1-1.png", + "caption": "Fig. 1 Inconel 718 forging Fig. 2 Turned and sliced testpieces", + "texts": [ + " This research investigates the evolution tool wear and machined surface roughness as a result of roughing and finishing during hole making operations. In this way the paper documents the evolution of tool wear expressed as VBmax (maximum value of tool wear on flank face) as well as evolution of machined surface roughness when drilling (roughing) and finishing. The cutting experiments were performed on Inconel 718 testpieces, made of a completely heattreated forging of a genuine turbine disk supplied by MTU-Aero Engines. Mechanical properties of Inconel 718 are shown in Table 1. The diameter of this forging was approx. 680 mm, see Fig. 1. Subsequently, this forging was turned to a flat disk with a uniform thickness of 22 mm, which afterwards was cut into segments by waterjet-cutting, see Fig. 2. The hardness of these segments was between 410 and 435 HV30. Fig. 3 shows one of the segments after removal from the machine tool after a series of drilling tests. The experimental drilling tests were performed on a five-axes machining center type MIKRON UCP 1050, equipped with SIEMENS Sinumerik 840D controller. The machine tool have the capability of applying coolant through an internal coolant supply (in the spindle), through external nozzles, and in both directions simultaneously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002042_tmag.2016.2527059-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002042_tmag.2016.2527059-Figure2-1.png", + "caption": "Fig. 2: Passive and bearing forces for a displaced rotor with zero angle estiamtion error (a) and non-zero angle estimation error (b).", + "texts": [ + " The force and torque calculation neglects z components of the flux due to the axial symmetry of the 0018-9464 (c) 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. bearingless disc drive. Therefore, it is sufficient to consider the radial and tangential flux components in the airgap. The calculation is carried out by integrating along the contour \u03be shown in Fig. 2. The coordinate \u03b1 depicts the position on \u03be. Force and torque are calculated as Fx \u221d \u222b 2\u03c0 0 ( B2 r \u2212B2 t ) cos (\u03b1)\u2212 (2BtBr) sin (\u03b1) d\u03b1 Fy \u221d \u222b 2\u03c0 0 ( B2 r \u2212B2 t ) sin (\u03b1) + (2BtBr) cos (\u03b1) d\u03b1 Tz \u221d \u222b 2\u03c0 0 BtBrd\u03b1. (1) The flux density components Br and Bt can be approximated by their space harmonics Bk,i as Bk (\u03b1) \u2248 \u2211 i B\u0302k,i sin (i\u03b1+ \u03c6k,i) , k \u2208 [r, t] . (2) The amplitude B\u0302k,i and phase \u03c6k,i of the space harmonics are dependent on the rotor position as well as the drive and bearing currents in the stator windings", + " The radial position controller of a bearingless machine controls the bearing currents in the stator windings to ensure a stable levitation of the rotor at a given reference position x\u2217, y\u2217. The controller requires information about the radial position and the angle of the rotor magnet to achieve this task. It is assumed that the radial position is known exactly but the rotor angle is only estimated. The actual rotor angle is denoted by \u03b8, the estimation by \u03b8\u0302 and the error by \u2206\u03b8 = \u03b8\u0302 \u2212 \u03b8. Assume that \u03b8 = 0, meaning that the rotor flux ~\u03a8R is placed on the x axis, there is no angle estimation error, \u03b8\u0302 = \u03b8, and the rotor is at a reference rotor position x\u2217 > 0, y = 0 as shown in Fig. 2(a). If higher flux density harmonics are neglected, the flux density components on the contour \u03be are Br,1 = B\u0302r,1,0 cos (\u03b1) Bt,1 = B\u0302r,1,0 cos (\u03b1\u2212 \u03c0/2) Br,2 = B\u0302r,2,\u2206 cos (2\u03b1) Bt,2 = B\u0302t,2,\u2206 cos (2\u03b1\u2212 \u03c0/2) (4) with B\u0302r,2 > B\u0302t,2. Inserting (4) into (1) shows that the radial displacement results in a force ~F\u2206. The bearing controller needs to generate a bearing force ~F \u2217 B that compensates ~F\u2206. Therefore, the radial position controller imposes bearing currents in the stator coils leading to the bearing flux density harmonics Br,2,B = B\u0302r,2,B cos (2\u03b1+ \u03c0) Bt,2,B = B\u0302t,2,\u2206 cos (2\u03b1+ \u03c0/2) , (5) which, accoring to (1), leads to a zero net force and zero torque. However, if the rotor angle estimate is not correct then the bearing will behave differently. Assuming, that the rotor angle is changed to \u03b8 > 0 as shown in Fig. 2(b). The passive flux density harmonics of order one and two are Br,1,0 = B\u0302r,1 cos (\u03b1\u2212 \u03b8) Bt,1,0 = B\u0302r,1 cos (\u03b1\u2212 \u03b8 \u2212 \u03c0/2) Br,2 = B\u0302r,2,\u2206 cos (2\u03b1\u2212 \u03b8) Bt,2 = B\u0302t,2,\u2206 cos (2\u03b1\u2212 \u03b8 \u2212 \u03c0/2) . (6) Assume that the radial position controller has no information about the new rotor angle, meaning that the estimated angle \u03b8\u0302 = 0\u21d4 \u2206\u03b8 > 0. Therefore, the bearing flux components are as given in (5). Inserting (5) and (6) into (1) results in FB,x < F \u2217 B,x, FB,y > F \u2217 B,y = 0, Tz < 0. (7) This shows that a rotor angle estimation error |\u2206|\u03b8 > 0 has two effects on the bearing behaviour", + " The transformed currents will lead to the forces Fx, Fy, according to (1), which define the rotor movement together with any external forces Fx,e, Fy,e. The external forces consist of the displacement force and possible disturbance forces. The observer block O calculates an estimate of the rotor angle error \u2206\u03b8\u0302 = \u2220 ( ~IB ) \u2212 \u2220 ( ~Fr,e ) \u2212 180\u25e6 (9) as the difference between the angle of the bearing current ~IB and the angle of an external radial force ~Fr,e. This force can be the gravitational force or a displacement force as shown in Fig. 2. The implementation of O is depending on the presence or absence of radial gravitational pull. If the external force is created by rotor displacement then the direction of the force can be calculated as \u2220~Fr,e = \u2220~r = arctan (y x ) (10) Therefore, the angle estimation error \u2206\u03b8\u0302 can be calculated by using ~IB and the rotor position x, y. Block A calculates \u03b8\u0302 based on \u03b8init and \u2206\u03b8\u0302 as \u03b8\u0302 (t) = \u03b8init + Pobs \u00b7\u2206\u03b8\u0302 (t) + Ii \u00b7 \u222b t 0 \u2206\u03b8\u0302 (\u03c4) d\u03c4 (11) with PA and IA being the proportional and integrator gain of the adaption mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003631_lra.2020.3033258-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003631_lra.2020.3033258-Figure1-1.png", + "caption": "Fig. 1. Our three-link robot experimental prototype.", + "texts": [ + " The purpose of this study is to model and analyze the dynamic bipedal frictional inchworm crawling locomotion, in order to comprehend and exploit the effects of inertia, for both soft and articulated robots. First, we develop a model for the hybrid dynamic [18] crawling (i.e. with discrete transitions between contact states) of a three-link robot with passive frictional contacts. The robot is actuated in open-loop by prescribing periodic inputs to its joints\u2019 angles. We study the effects of frequency and other input parameters on the crawling gait, and find trends and optimal performance. We manufacture and experimentally test a three-link robot (Fig. 1) with good quantitative and 2377-3766 \u00a9 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 22,2020 at 01:29:15 UTC from IEEE Xplore. Restrictions apply. excellent qualitative agreement with the theoretical predictions, which proves the applicability of our analysis (see supplementary video [19]). We also investigate the effects of friction uncertainties, which was shown to have major influence [20], and mass asymmetry, propose a novel input shaping technique and apply machine-learning based optimization to improve the performancein traveling distance. Finally, we discuss a feedback control strategy. In order to investigate crawling at frequency range where inertial effects are significant, we have manufactured the three-link robot prototype in Fig. 1. The robot has a central link with length l0, mass m0 and moment of inertia J0 and two distal links, with length li, mass mi and moment of inertia Ji, for i = 1, 2 (see Fig. 2). For most of this work we consider identical distal links l1 = l2 \u2261 l, m1 = m2 \u2261 m, J1 = J2 \u2261 J , and in Section VII we investigate the influence of asymmetric mass distribution. The experimental setup parameters are summarized in Table I. Two servomotors at the joints receive a sequence of angle commands from the microcontroller (in open-loop) and track it with internal closed-loop control", + " \u201cflatter\u201d robot), by varying the input frequency \u03c9 a direction reversal occurs, as shown in Fig. 7(a) in solid blue curve (for \u03d50 = 145\u25e6, Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 22,2020 at 01:29:15 UTC from IEEE Xplore. Restrictions apply. A = 18\u25e6 and \u03c8 = 20\u25e6). The same phenomenon is observed for a sweep of the nominal angle \u03d50 at a frequency in the above range (\u03c9 = 9 [rad/s]) in Fig. 7(b). In order to measure the performance of our robot in Fig. 1, we video it with a simple webcam. The video lens distortion is then corrected with MATLAB Computer Vision System Toolbox and post-processed with Kinovea software, which gives planar x-y position of selected points on the robot\u2019s links. Since the robot is controlled in open-loop, as the frequency increases the servo struggles to follow the prescribed amplitude. Therefore we initially calibrate the amplitudes of inputs\u2019 reference trajectories to achieve the desired angles in (5), and validate the other parameters (such as frequency)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure4-1.png", + "caption": "Figure 4. Crawler shoe link pin joints.", + "texts": [ + " D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 Ku\u20acu \u00bc E (6) E \u00bc Ku _u\u00f0 \u00deu _u 2Kut _u Ktt (7) _ui \u00bc \u00bd _Bxi _Byi _Bzi _wi _hi _/i T vector of generalised velocity of body i (8) \u20acui \u00bc \u00bd\u20acBxi \u20acByi \u20acBzi \u20acwi \u20achi \u20ac/i T vector of generalised acceleration of body i (9) Ku \u00bc @K @u Jacobian of the constraint equation \u00f04\u00de (10) Kt and Ktt are the partial derivatives of constraint Equation (4) with respect to t; Kut is the partial differential of Equation (10) with respect to t. The algebraic joint and motion constraint equations for the crawler track assembly and flat terrain shown in Figure 3 are described for the following constraints. In the actual crawler track assembly, the crawler shoes are connected to each other using two link pins as shown in Figure 4. The link pins are made to rotate about the axis of the rotation as shown in Figure 4. This rotation action reduces loading in the shoe lugs and increases the life of the link pin [3]. It can be seen from Figure 4 that each crawler shoes can have only relative rotational motion about the joint y-axis due to the pin joint. The relative rotation about the joint z and x axes, and the relative translation motion along the joint x, y and z axes are constrained and negligible. This pin joint has one rotational DOF and is called a revolute joint [15, 17]. Many researchers [4\u20136, 12, 18\u201322] used revolute or pin joints for their analysis. In a majority of these studies, the connection between the track links for slow moving tracked vehicle was simulated using revolute joints with one DOF", + " Rubinstein and Hitron [6] and Madsen [18] added friction to the revolute joint to simulate the contact forces between pin and hole. Since the link pin connecting two crawler shoes undergoes rotation within the lugs, the effect of friction can be neglected. The clearance between the link pin and the lug hole is very small compared to the dimensions of the lugs and can be neglected. Therefore, this study models the crawler shoe links as frictionless revolute joint with one DOF. The axis of two link pin joints is collinear as shown in Figure 4 based on the kinematic constraint equations for the crawler tracks. This results in redundant constraints when revolute joints are used for two link pins to connect two rigid crawler shoes. To D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 remove the redundant constraints, one pin joint is made a spherical joint and another pin joint is made the parallel primitive joint as shown in Figure 4 [13,14]. The spherical joint removes three relative translations along the joint x, y and z axes, while the parallel primitive joint removes two relative rotations about the joint z and x axes. The combination of both spherical and parallel joints results in an equivalent revolute joint between two crawler shoes. The bodies, labelled 2\u201314, are connected by revolute joints at the locations of the first link pin. In Figure 5, the points Pi and Pj defined along the joint axis on bodies i and j are always coincident during the entire motion of the crawler track [15]", + " KS ui; uj \u00bc rPi rPj \u00bc Bi \u00fe Ais 0 Pi Bj Ajs 0 Pj \u00bc 0 (14) In Equation (14), Ai and Aj are the rotation matrices that define the orientation of the body i and j with respect to the global coordinate system [15], and s0Pi and s0Pj are the positions of the point P on body i and body j with respect to its centroidal coordinate system. For the body i, Ai and s0Pi can be defined as Equations (15) and (16). Ai \u00bc cos/i coswi coshi sinwi sin/i sin/i coswi coshi sinwi cos/i sinhi sinwi cos/i sinwi\u00fe coshi coswi sin/i sin/i sinwi\u00fe coshi coswi cos/i sinhi coswi sinhi sin/i sinhi cos/i coshi 2 4 3 5 (15) s0Pi \u00bc s0x;Pi s0y;Pi s0z;Pi h iT (16) D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 The bodies, labelled 2\u201314, are connected by parallel primitive joint at the second link pin location shown in Figure 4. In Figure 6, the vectors ai, bi and ci and vectors aj, bj and cj are defined along the joint x, y and z coordinate axis defined on bodies i and j at the joint location point P. [13,15]. The two constraint equations for parallel primitive joints are from Equation (17). KP ui; uj \u00bc ci bj ai bj \u00bc 0 (17) i = 2, 3, \u2026, 13 and j = 3, 4, \u2026, 14 The propel action of the crawler track assembly is studied for two types of prescribed motion specified at the centre of mass of Part 14 (crawler shoe 13) (point P14) as shown in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003961_j.measurement.2020.108909-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003961_j.measurement.2020.108909-Figure4-1.png", + "caption": "Fig. 4. Hexapod Felix.", + "texts": [ + " 1 shows, the optical sensor head scans the scale tape and builds an incremental measuring system with a resolution of 0.1 \u03bcm. We experimentally validated in our work [23] that the IMD DBB could reach a comparable accuracy (1 \u03bcm in 5 mm measuring range) to the commercial DBB (\u2264 1 \u03bcm in \u00b10.1 mm measuring range by QC20W [10]). Fig. 2 visualizes the physical object of IMD DBB, the material of which is under 1000 euro cost range. Before each measurement, the DBB would be calibrated with an external calibrator (Fig. 3) and the absolute bar length refers to the calibration value. We chose Felix (Fig. 4) as the target platform for the experimental validation: a Stewart\u2013Gough platform4 with a simple 6\u20136 structure, 3 We will occasionally use the term \u2018spatial\u2019 to denote 6-D measurement, which includes not only the end effector/spindle\u2019s position in 3-D space but also the orientation thereof. 4 Although not exactly equivalent, it is also occasionally called hexapod in the literature. In this article, we will use both terms interchangeably. which was also designed and constructed at IMD. The active translational joints of Felix consist of 6 identical electric cylinders and are controlled via Beckhoff TwinCAT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001229_acc.2013.6580137-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001229_acc.2013.6580137-Figure1-1.png", + "caption": "Fig. 1. Desired formation of the set of eight agents and a leader with information flow topology.", + "texts": [ + " The dynamical equations of motion for the ith agent are given by q\u0308i(t) = ui(t) + gi(qi(t), q\u0307i(t)), qi(0) = q0i, q\u0307i(0) = q\u03070i, t \u2265 0, i = 1, ...,m, (59) where qi \u2208 R n is the generalized position vector for the ith agent, ui(t) \u2208 R n, t \u2265 0, is the control input for the ith agent, and gi(qi, q\u0307i) \u2208 R n represents bounded uncertainties and disturbances affecting the dynamics of the ith agent. For the following simulation, we consider a case of planar coordinated motion of eight double integrators and a leader. The desired formation and the information flow between agents are shown in Figure 1. This communication topology remains unchanged over time. Note that the directed graph of the formation satisfies the required property that there exits at least one directed path from the leader to any agent. Also, note that L2 \u2208 R 8\u00d78 given by (13) is invertible. Next, the leader is set to be moving counter-clockwise on a circular path of radius 1 according to xL(t) = cos(\u03c0t/10), yL(t) = sin(\u03c0t/10), t \u2265 0. The values of control gains are set to be Ki = 2 \u00d7 I2 and Kiaux = 2 \u00d7 I2, i = 1, . . . , 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001870_iccas.2014.6987548-FigureI-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001870_iccas.2014.6987548-FigureI-1.png", + "caption": "Fig. I.Rotary Inverted Pendulum", + "texts": [ + " 8 and the input voltage needed for stabilization IS illustrated in Fig. 9. The vibrations of the voltage are high due to the switching effect of the sliding mode controller. For the practical implementation, the Quanser Rotary Inverted Pendulum illustrated in Fig. 10 was used. LabView helped to create the graphical user interface of the RIP. The optical encoder of the pendulum considers 180 degrees as the position of stability, which represents the same reality with the simulation requirement under some reference transformation. It can be seen from Fig II. that both arm and pendulum are stablized. For the energy that was used to swing up the pendulum and to stabilize is shown in Fig. 12. The controller has contributed negatively with vibrations. The umneasured states of rotary inverted pendulum were extracted by a higher order differentiator and the highly nonlinear system was stabilized by sliding mode controller that was designed. [t is proposed that another sliding mode strategy for this model of the pendulum to avoid stronger vibrations in future work and an adaptive model of that control strategy be investigated in future work" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000760_peds.2013.6527174-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000760_peds.2013.6527174-Figure13-1.png", + "caption": "Fig. 13. Construction of Shield structure DrSRM.", + "texts": [ + " However, it was difficult. So, we propose a shield structure to improve the characteristics. x y z x y z (a) (b) Fig. 10 Analysis model of DrSRM (a) 2-D, (b) 3-D. 0 10 20 30 40 50 60 70 0 2 4 6 8 10 A v er ag e to rq ue [ N -m ] Current [A] 2-D FEM 3-D FEM Fig. 11. Average torque-Current for 2-D FEM and 3-D FEM. 0 50 100 150 200 250 300 350 400 0 3 6 9 12 15 2D_1A 2D_5A 2D_10A 3D_1A 3D_5A 3D_10A In d u c ta n c e [ m H ] Rotational angle [deg] Fig. 12. Inductance waveform. IV. SHIELD STRUCTURE Fig. 13 shows shield structure of the DrSRM. As shown in this figure, a high magnetic permeability material is affixed to both side surfaces of the stator yoke. This material is permendur Fe-Co alloy with a maximum flux density within the soft magnetic material. Fig. 14 shows the characteristics of a permendur and usual electric steel 50H1300. 0 0.5 1 1.5 2 2.5 0 1000 2000 3000 4000 5000 6000 7000 Permendur 50H1300 M ag n et ic f lu x d en si ty [ T ] Magnetic field strength [A/m] Fig.14 . Magnetic characteristics for permendur and 50H1300" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003903_cdc42340.2020.9304420-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003903_cdc42340.2020.9304420-Figure4-1.png", + "caption": "Fig. 4: Two networks and their leaders (gray). In (a): \u03b4(G, dset(G,V`)) = 5, \u03b4(G,V`) = 4, \u03b6(G,V`) = 3. In (b): \u03b4(G, dset(G,V`)) = 9, \u03b4(G,V`) = 6, \u03b6(G,V`) = 5.", + "texts": [ + " Hence, the longest possible PMI sequence of DL vectors with the additional leaders can not be shorter, i.e., \u03b4(G, dset(G,V`)) \u2265 \u03b4(G,V`). Remark 4.3: While Theorem 4.2 shows that the combined bound is at least as good as the distance-based and zero-forcing-based bounds, it should also be emphasized that there exist networks G = (V,E) and leader sets V` \u2286 V , where the combined bound is strictly better than the two original bounds, i.e., \u03b4(G, dset(G,V`)) > \u03b4(G,V`), \u03b6(G,V`). We provide two such examples in Fig. 4. We compare the lower bounds on the dimension of strong structurally controllable subspace on Erdo\u0308s-Re\u0301nyi (ER) and Baraba\u0301si-Albert (BA) graphs. ER graphs are the ones in which any two nodes are adjacent with a probability p. BA graphs are obtained by adding nodes to an existing graph one at a time. Each new node is adjacent to \u03b5 existing nodes that are chosen with probabilities proportional to their degrees. In all the simulations, we consider undirected graphs with n = 100 nodes. In Figs. 5 and 6, we plot lower bounds on the dimension of SSCS, including \u03b4(G,V`), \u03b6(G,V`) and \u03b4(G, dset(G,V`)), as a function of number of leaders |V`| = `" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001479_ijma.2014.064096-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001479_ijma.2014.064096-Figure4-1.png", + "caption": "Figure 4 Kinematics of a four-axis Cartesian robot with three translational (q1, q2 and q3) belt drive axes and one rotational axis (q4)", + "texts": [ + " 5 Compose the parameter vector \u03c1 from the variable parameters {ts,i, Ti} (compare Section 3.1). 6 Set the linear inequality constraints (A\u03c1 \u2264 b), see (21). 7 Apply global optimisation algorithm (e.g., PSO, GA). 8 Run optimisation process. Exemplary trajectory optimisation results from simulations are presented in Section 4 that illustrate the amplification of energy exchange and the resulting reduction of energy demand. To demonstrate the principle of the presented trajectory optimisation approach, the procedure has been carried out using the model of a Cartesian robot (see Figure 4). The robot kinematics include four axes, actuated with servo drives, that are electrically coupled using a DC bus. In the chosen multi-axis system, no mutual mechanical interferences appear within the axes. Therefore, no mechanical but electrical energy is transferred between the servo drives to emphasise the desired effect of the coupled DC bus. The model parameters have been collected from data sheets of belt drive and servo drive manufacturers. A load of 10 kg is attached to the EE and the basic construction parameters are given in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure7.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure7.3-1.png", + "caption": "Figure 7.3 D\u2013H convention for the first joint.", + "texts": [ + "16) \u2217 The offset between the unit vectors u\u20d7(j) 1 = u\u20d7(jk) 1 and u\u20d7(k) 1 = u\u20d7(kj) 1 : sk = sjk = OjkOkj = BkOk along u\u20d7(k) 3 (7.17) \u2217 The effective link length of the link j between the joints ji and jk : bj = bik = bijk = OjiOjk = OjBk along u\u20d7(j) 1 (7.18) \u2217 The joint frames ji(Oji) and jk(Ojk) of the kinematic elements ji and jk on the link j: The joint frame ji(Oji) and the link frame j(Oj) are taken to be coincident. That is, ji(Oji) = j(Oj); u\u20d7(ji) k = u\u20d7(j) k for k = 1, 2, 3 (7.19) jk(Ojk) = jk(Bk); u\u20d7(jk) 1 = u\u20d7(ji) 1 = u\u20d7(j) 1 , u\u20d7 (jk) 3 = u\u20d7(kj) 3 = u\u20d7(k) 3 (7.20) Figure 7.3 shows a sketch that involves the base link (0) and the first link (1) of a serial manipulator. In general, the base frame 0(O0) is attached to 0 independently of how the manipulator is installed. However, the joint frame 01(O01) is attached to 0 in accordance with the first axis of the manipulator. So, 01(O01) is also called a manipulator-specific base frame. The third and first basis vectors of 01(O01) are such that u\u20d7(01) 3 = u\u20d7(1) 3 and u\u20d7(01) 1 is along the common normal 01 between the axes along u\u20d7(0) 3 and u\u20d7(1) 3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003047_icit45562.2020.9067209-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003047_icit45562.2020.9067209-Figure9-1.png", + "caption": "Fig. 9. Time Harmonic simulation to get the rotor response currents to the stator slot harmonics.", + "texts": [ + " Often the the rotor bars are skewed to reduce the torque ripple and the cogging effects. However these effects can be reduce by choosing a correct combination of stator and rotor slot number. Further, the common skew angle is the stator slot angle. Thus theoretically, the rotor winding has a winding factor equal to zero for these harmonics: it is not able to reply to them. This can be proved even using finite elements. The rotor induced current by the slot harmonics is computed by means of TH simulations. An example of a TH simulation is reported in Fig. 9. In presence of skewing a series of TH simulations is done, considering several slices of the machine with different displacement, to get the average value of each bar current, considering the skewing [13]. The comparison between the rotor induced currents with and without skewing is reported in Fig. 10. It is clear that, in presence of the skewing of one stator slot angle, the rotor reaction to the slot harmonics is negligible. As a manner of facts, in this case, a reduction of the rotor iron additional losses is not possible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003932_s40962-020-00549-5-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003932_s40962-020-00549-5-Figure2-1.png", + "caption": "Figure 2. Manufacturing process of enclosed wheel: (a) manufacturing the plaster core molds using FDM equipment, (b) manufacturing the plaster cores, (c) assembling the plaster cores, (d) casting, (e) finish.", + "texts": [ + " Those studies16,17 focused on the surface finishing and dimensional accuracy via different process parameters and defined the main factors that affect the dimensional accuracy are layer thickness, build orientation and filled density. Liu Xinhua et al.18 constructed the theory model of the process of distortion of thin plate made by PLA and determined the reasonable process factors to minimize its distortion through actual experiments. In the present work, we have designed the optimal profile of blade grooves using finite element analysis and established the process of manufacturing the wheel by the plaster casting process using FDM (Figure 2). As shown in Figure 2, we make the plaster mold; its material is PLA, using FDM machine. And then, we manufacture the plaster cores by the plaster mold setup and assemble the plaster cores in the casting mold. Lastly, molten aluminum alloy is poured into the casting mold and the turbo-expander wheel is finished through the machine process. The wheel\u2019s 3D model which was designed with SolidWorks is shown in Figure 3. International Journal of Metalcasting There are 20 blade grooves in the wheel and seal thresholds for air tightness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002593_978-3-319-46155-7_6-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002593_978-3-319-46155-7_6-Figure12-1.png", + "caption": "Fig. 12 A new way of designing functional parts by the integration of lattice structures", + "texts": [ + " 2010). Nevertheless, there is no comprehensive collection of mechanical properties of lattice structures under compressive, tensile, shear and dynamic load. The deformation and failure mechanisms are not studied sufficiently. A relatively new field of research is the influence of different scan parameters/strategies on mechanical properties. To reach the overall objective of our research, these challenges need to be overcome to design functional adapted parts with integrated lattice structures (see Fig. 12). Economist (2011) Cover story. 2 Dec 2011 EOS GmbH (2014) Industry: Siemens\u2014EOS Technology opens up new opportunities for indus- trial gas turbine maintenance cost reduction. Available via http://www.eos.info/press/cus tomer_case_studies/siemens. Accessed 23 Jun 2016 Gibson I, Rosen D, Stucker B (2010) Additive manufacturing technologies. Rapid prototyping to direct digital manufacturing. Springer, Heidelberg GE Report (2014) Fit to print: new plant will assemble world\u2019s first passenger jet engine with 3D printed fuel nozzles, next-gen materials" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001887_0959651816629540-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001887_0959651816629540-Figure1-1.png", + "caption": "Figure 1. Missile configuration.", + "texts": [ + " Possible option is propulsion-based control in the form of a reaction jet control system (RCS) combined with fins\u2019 control,3\u201319 and the compensated missile is often called dual-control missile.16\u201319 There are two types of RCS configuration, the direct force type and the moment type. The RCS is located near the center of gravity (c.g.) arising the acceleration directly in the former type, and in ahead of or in rear of c.g. arising the force and the moment in the later type.5 The article here concentrates on the moment-type missiles, which is based on PAC-3 configuration1 as shown in Figure 1, where 180 reaction jets of 10 circles are located in radical direction evenly at the distance of 1m before the center of mass. Reaction jets could provide a huge moment in a short time and would not be affected by the atmosphere around, thus providing the missile excellent maneuverability in the upper air.1 However, the participation of the reaction jets brings coordination problem with fins. Moreover, it brings intensive disturbances to the missile aerodynamics. Besides, the anti-air missile dynamics undergoes substantial change during its flight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003895_cdc42340.2020.9304246-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003895_cdc42340.2020.9304246-Figure2-1.png", + "caption": "Fig. 2. Analysis of the condition for satisfying p\u2217di = p\u2217i . When |\u03c6i| \u2264 \u03c0/2 (left), p\u2217di = p\u2217i is always satisfied because the gray hatched area includes the blue hatched area. However, when |\u03c6i| > \u03c0/2 (right), p\u2217di = p\u2217i is not always satisfied because the gray hatched area only partially includes the blue hatched area.", + "texts": [ + " Although the stationary positions p\u2217d \u2208 R2N do not guarantee local optimality of problem (15), we show the necessary and sufficient condition for satisfying (\u2202C\u2217d)N \u2286 (\u2202C\u2217)N . For tractability, we introduce the following additional assumption. Assumption 4: The object C is convex. From Assumption 1, the boundary of the object is left/righthand differentiable. We here introduce two functions cil : R \u2192 R and cir : R \u2192 R, which represent left/righthand side of local boundary at pi in local coordinate (xi, yi) as shown in Fig. 2. We consider two adjacent agents mod(i\u2212 1, N) and mod(i+1, N) of i, where mod(\u2217, \u2217) represents the modulo operation. Then, from Assumption 4, the position pi is a local optimal position p\u2217i \u2208 R2 for all i (i = 1, 2, . . . , N) if and only if there exist ali, ari \u2265 0 such that pi \u2212 1 N N\u2211 j=1 pj = ali [ \u2207\u2212cil(pix) \u22121 ] + ari [ \u2207+cir(pix) \u22121 ] i = 1, 2, . . . , N (26) where \u2207\u2212 (\u2207+) represents left (right)-hand derivative. Meanwhile, from Assumption 4, the position pi is a stationary position p\u2217di \u2208 R2 for all i (i = 1, 2, ", + " Proof : From Lemma 2, the dynamics (20) converges to the set (\u2202C\u2217d)N of stationary positions p\u2217d \u2208 R2N . Since all elements of (\u2202C\u2217d)N satisfy (26), we obtain (\u2202C\u2217d)N \u2286 (\u2202C\u2217)N . The following discussion on the object shape would be beneficial to judge (\u2202C\u2217d)N \u2286 (\u2202C\u2217)N is satisfied or not. From Assumption 4, there exist dli, dri \u2265 0 such that pi \u2212 1 N N\u2211 j=1 pj = dli(pi \u2212 pmod(i\u22121,N)) + dri(pi \u2212 pmod(i+1,N)). Hence, when the magnitude of the object angle |\u03c6i| at pi is less than or equal to \u03c0/2, p\u2217di = p\u2217i is always satisfied as illustrated in Fig. 2 (left). However, when |\u03c6i| > \u03c0/2, p\u2217di = p\u2217i is not always satisfied as illustrated in Fig. 2 (right). Thus, only when |\u03c6i| > \u03c0/2, we need to check whether p\u2217di satisfies (26). We next discuss the combined optimization dynamics (14) and (20). To this end, we formulate an augmented error system \u03a3\u03be of systems \u03a3u, \u03a3\u00b5 and \u03a3x regarding the equilibrium point (u\u2217(p\u2217), \u03b2\u2212 1 2\u00b5\u2217(p\u2217), \u221a \u03b1\u03b2 1 2xref) as follows: \u03a3\u03be : \u03be\u0307 = \u2212I2N \u2212B>\u03b2B \u2212B>\u03b2 1 2 \u2212 \u221a \u03b1B>\u03b2 1 2 \u03b2 1 2B 0 I3\u221a \u03b1\u03b2 1 2B 0 0 \u03be + B>\u03b2\u03b2 1 2 \u221a \u03b1\u03b2 1 2 f(x) + 0 0[ \u2212 \u221a \u03b1mR\u0307>ow\u03bd w ref 0 ] := A\u03be + g(x) where we have defined \u03be := [ u> \u00b5>\u03b2\u2212 1 2 \u221a \u03b1x>\u03b2 1 2 ]> \u2212[ (u\u2217(p\u2217))> (\u00b5\u2217(p\u2217))>\u03b2\u2212 1 2 \u221a \u03b1x>ref\u03b2 1 2 ]> " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003799_s40313-020-00664-y-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003799_s40313-020-00664-y-Figure1-1.png", + "caption": "Fig. 1 Dump truck roll model", + "texts": [ + " Secondly, the sprung mass except the cargo box and the cargo is equivalent to the frame sprung mass, and the part is regarded as a rigid body, and the flexible cargo box is simplified into a mass and a pair of spring dampers, and the distance between the two spring dampers is equal to the distance between the flip hinges. Thirdly, the lateral deformation of the tire, and all tires do not slip laterally on the road surface are ignored. The roll model of the dump truck during the lifting of the cargo box is shown in Fig. 1. Where m1 is the unsprung mass of the vehicle; m2 is the sprung mass of the frame; m3 is the mass of the cargo box and cargo; A1, A2 and A3 are the roll centers of m1, m2 and m3; \u03d5u , \u03d5s1 and \u03d5s2 are the roll angles of m1, m2 and m3; G1, G2 and G3 are the gravity of each part of the vehicle; \u03b2 is the lateral slope angle of pavement; Y0 is the initial lateral offset of the centroid of m3; d is the wheelbase; hu is the roll length of m1; hr1 is the initial height of the roll center A2; ul and ur are the active suspensions control forces; Fl and Fr are vertical loads acting on the left and right wheels; kt is the vertical stiffness of a single tire; k1 and k2 are the stiffness of the suspension spring and the equivalent stiffness of the box when lifting; c1 and c2 are the equivalent damping coefficients of the suspension and cargo" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003443_j.procir.2020.04.140-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003443_j.procir.2020.04.140-Figure6-1.png", + "caption": "Fig. 6. Clearance vector loop model of a 4-bar-mechanism with joint clearance", + "texts": [ + " Furthermore, lubricants in the non-assembly mechanism are not existing, so a contact load can occur, whereby the resulting impulse forces are transferred to following mechanical parts. These impulses and the subsequent continuous contact can be modeled by the force model [8]. According to Flores et al. [13], the size of the clearance is hereby defined by the difference in radius between bearing Rb and journal Rj: c = Rb \u2212 Rj. (5) The vector loop approach for the 4-bar mechanism of Fig. 3, is now extended by the clearance vectors c12 and c23 for the clearance affected joints J12 and J23 in Fig 6. According to Goessner (see Eq. 1), for a 4-bar mechanism consisting of four links g and four joints n, one vector loop is sufficient for the characterization of its motion behavior, whereby the informations for the clearance vectors are derived through the MBS [23]: L1 \u00b7 e j\u00b7\u03b81 + c12 \u00b7 e j\u00b7\u03b312 + L2 \u00b7 e j\u00b7\u03b82+ c23 \u00b7 e j\u00b7\u03b323 \u2212 LAB \u2212 L3 \u00b7 e j\u00b7\u03b83 = 0. (6) 3.3. Statistical tolerance analysis using sampling techniques Statistical tolerance analysis using sampling techniques offers the possibility to determine the influence of joint clearances and geometrical deviations resulting from the AM-process on the movement behavior", + " Furthermore, lubricants in the non-assembly mechanism are not existing, so a contact load can occur, whereby the resulting impulse forces are transferred to following mechanical parts. These impulses and the subsequent continuous contact can be modeled by the force model [8]. According to Flores et al. [13], the size of the clearance is hereby defined by the difference in radius between bearing Rb and journal Rj: c = Rb \u2212 Rj. (5) The vector loop approach for the 4-bar mechanism of Fig. 3, is now extended by the clearance vectors c12 and c23 for the clearance affected joints J12 and J23 in Fig 6. Fig. 6. Clearance vector loop model of a 4-bar-mechanism with joint clearance According to Goessner (see Eq. 1), for a 4-bar mechanism consisting of four links g and four joints n, one vector loop is sufficient for the characterization of its motion behavior, whereby the informations for the clearance vectors are derived through the MBS [23]: L1 \u00b7 e j\u00b7\u03b81 + c12 \u00b7 e j\u00b7\u03b312 + L2 \u00b7 e j\u00b7\u03b82+ c23 \u00b7 e j\u00b7\u03b323 \u2212 LAB \u2212 L3 \u00b7 e j\u00b7\u03b83 = 0. (6) 3.3. Statistical tolerance analysis using sampling techniques Statistical tolerance analysis using sampling techniques offers the possibility to determine the influence of joint clearances and geometrical deviations resulting from the AM-process on the movement behavior" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003466_s106836662004008x-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003466_s106836662004008x-Figure3-1.png", + "caption": "Fig. 3. A general view of the friction lining 1 with block 2.", + "texts": [ + " 4 2020 high coefficient of friction and wear resistance and is widely used in brake and friction units of machines and mechanisms for dry friction with the surface friction temperature up to 180\u2013200\u00b0C and contact pressure up to 1.5 MPa [6]. The cross section of the friction lining is a square with a side of 10 mm and thickness of 8 mm. The friction lining is placed in a special block, which is fixed on the lever that creates a load on the friction pair with the help of measured loads of various weights. A general view of the friction lining with the block is shown in Fig. 3. The rotating disk is made of steel S235JR (the Russian analog is steel St3), the average friction radius is 20 mm, and the disk thickness is 10 mm. The studies were carried out in the stationary friction mode at the ambient temperature of 20\u201325\u00b0C with a wide range of loads (Table 1). The time of each test was chosen so that the friction pair for at least 50% of the test time was working in steady state thermal conditions. Therefore, the duration of each test was 2 h, after which the T-11 friction machine was automatically turned off" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003008_j.aej.2020.03.015-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003008_j.aej.2020.03.015-Figure1-1.png", + "caption": "Fig. 1 Schematic view of four pads tilting pad jour", + "texts": [ + " From the results for both approaches it can be outlined that neural network could model bearing systems in real-time applications. Sinanoglue et al. [11] presented and analyzed the pressure variation on steel shaft of plain journal bearings experimentally with different speeds and temperature conditions and applied a feed forward neural network to predict the bearing performance. A tilting pad journal bearing (TPJB) is a type of fluid film bearing that supports a rotating machinery. It has the advantage of minimizing the destabilizing forces and reduced loss of shear power compared to a plain journal bearing. Fig. 1, shows the geometry of the bearing and the coordinate system. A tilting pad journal bearings normally has three to six pads. Each pad supported by a pivot and tilts about it during operation and forms a convergent wedge between the inner surface of the pad and the shaft. The journal shear drives a viscous fluid film into the wedge to generate a hydrodynamic pressure field allowing the bearing to carry the applied load (W). The load carrying capacity of a journal bearing depends on the pressure produced in the lubricant layer during journal rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002949_s12541-019-00277-9-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002949_s12541-019-00277-9-Figure2-1.png", + "caption": "Fig. 2 Drive-side mesh stiffness calculation model in any transverse plane", + "texts": [ + " Moreover, as the NTCIG needs to be reversed frequently in locating and tracking precision transmission systems, this paper places significant emphasis on the twosided tooth mesh analysis through the DTE of the gear pair based on the angular position detection of the anti-backlash strategy under varying load and speed excitations. The NTCIG tooth thickness varies due to differences in the tangential modifications and similarities in the radial modifications, as shown in Fig.\u00a01. Compared with the conventional variable tooth thickness gear, each transverse plane of the NTCIG has identical dedendum and addendum circles. Therefore, gear interference can be avoided during the axial movement of the driven gear to eliminate backlash. The coordinate system is built based on the rotation center of the driving gear. In Fig.\u00a02, the rotational angle \u03b8 of the driving gear is defined as the angle between the symmetrical line of the driving gear tooth and the horizontal plane. The NTCIG is divided into N equal pieces along the direction of the gear width b. When N is sufficiently large, the drive-side mesh stiffness of each piece is calculated as the involute spur gear with constant tangential and radial modifications. Therefore, the gear tooth parameters for any piece along the gear width direction illustrated in Fig.\u00a02 can be obtained as follows: (1) Sk = 2rk sin ( + 4x tan 0 2z + inv 0 \u2212 inv k \u00b1 ||bi||tan r ) (2) k = arccos(rg\u2215rk) (3)SF = \u23a7 \u23aa\u23aa\u23aa\u23a8\u23aa\u23aa\u23aa\u23a9 2rF sin \ufffd \ud835\udf0b+4x tan \ud835\udefc 0 2z + inv\ud835\udefc 0 \u2212 inv\ud835\udefcF \u00b1 \ufffdbi\ufffdtan\ud835\udefd r \ufffd if rb \u2264 rF 2rg sin \ufffd \ud835\udf0b+4x tan \ud835\udefc 0 2z + inv\ud835\udefc 0 \u00b1 \ufffdbi\ufffdtan\ud835\udefd r \ufffd if rb > rF 1 3 In Eqs.\u00a0(1) and (3), the + signifies an increase in the tooth thickness relative to the middle transverse plane, and \u2212 signifies the opposite. For any piece, the total deformation \u03b4i of the tooth along the transverse plane under the mesh force on the action line can be calculated as: Each deformation can be calculated from the gear tooth parameters using Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003560_icra40945.2020.9197167-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003560_icra40945.2020.9197167-Figure5-1.png", + "caption": "Fig. 5: Adding visibility nodes. If connections between test points ( ) are blocked and the node is visible to the current robot position, it is added to the sparse graph with a connection to the goal node. Additionally, a position node is added at the robot\u2019s current location to maintain graph structure. Newly added edges are shown as dashed purple lines.", + "texts": [ + " Points along the straight-line connections between the test node and its neighbors are checked for collision with the kd-tree local map. If more than one connection is blocked and the test node is visible from the current robot position, it is added to the sparse graph as a visibility node with an edge connection to the goal node ng . When a visibility node is added, a position node at the robot\u2019s current location is also incorporated into the graph to maintain structure. The process for adding a visibility node is shown in Fig. 5. To add a node into the sparse graph, three quantities are needed: the id of the node(s) to which it connects, the Euclidean distance (cost) between the nodes, and whether the edge is traversable, i.e., has the robot physically moved between the two nodes. An example of a non-traversable connection is the one made between a visibility node and the goal node. The edge cost for a visibility node is the Euclidean distance estimate cost-to-goal from the visibility node. A series of adjacency vectors store the sparse graph to allow for efficient modification and search" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002949_s12541-019-00277-9-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002949_s12541-019-00277-9-Figure6-1.png", + "caption": "Fig. 6 Additional angle for the anti-backlash NTCIG", + "texts": [ + " Figure\u00a05 shows that the results of the NTCIG are worse than the results of the spur gear, which demonstrates that the gear pair mesh stiffness decreased because of tangential modifications. However, the mesh stiffness of the NTCIG varies continuously because of the continuous changes in the tooth contact line, which slows the abrupt changes in the mesh stiffness of the spur gear pair. Moreover, both the drive-side and the back-side mesh stiffnesses for the NTCIG and the spur gear pair are symmetric with respect to the line of x = 0 because of the gear teeth symmetry. 3 NLTE Calculation of\u00a0NTCIG with\u00a0Anti\u2011backlash Figure\u00a06 illustrates the anti-backlash principle of the NTCIG in that the meshing tooth thickness of the gear pair on every transverse plane changes when the driven gear moves along the axial direction. The relationship between the axial displacement (s) and the backlash (Bh) can be expressed as follows: (10)kb( ) = kd(\u2212 ) 1 3 Based on the displacement increment along the line of action shown in Fig.\u00a07, the detailed computations of NLTEd, NLTEb, and Bh for the NTCIG with eccentricity and center distance errors can be expressed as: [1, 21] 4 DTE Calculation of\u00a0NTCIG with\u00a0Anti\u2011backlash Figure\u00a08 illustrates the dynamic model of the NTCIG with a dynamic equation that is expressed as [23]: (12) NLTEdanti = NLTEd + s tan R 2 NLTEbanti = NLTEb + Bh \u2212 s tan R 2 (13) \ufffd \u0394F 1d \u0394F 2d \ufffd = \ufffd e1 0 0 e2 \ufffd\ufffd sin( 1 + + 10 ) \u2212 sin( + 10 ) sin( 2 \u2212 + 20 ) \u2212 sin(\u2212 + 20 ) \ufffd \ufffd \u0394F 1b \u0394F 2b \ufffd = \ufffd e1 0 0 e2 \ufffd\ufffd sin( 1 \u2212 + 10 ) \u2212 sin(\u2212 + 10 ) sin( 2 + + 20 ) \u2212 sin( + 20 ) \ufffd NLTEd = 180 \u2217 60 R 2 cos \u22c5 2\ufffd i=1 \u0394Fid NLTEb = 180 \u2217 60 R 2 cos \u22c5 2\ufffd i=1 \u0394Fib Bh = 180 \u00d7 2 sin R 2 cos \ufffd e 1 e 2 \u0394a \ufffd\u23a1\u23a2\u23a2\u23a2\u23a3 \u2212 cos( 1 + 10 ) cos( 2 + 20 ) 2 sin \u23a4 \u23a5\u23a5\u23a5\u23a6 In the anti-backlash process, the driven gear moves towards the dotted line position from the solid line position so that an additional angle is produced, as shown in Fig.\u00a06. Therefore, the drive- and back-side no-load transmission error (NLTE) during the anti-backlash process is modified as: (11)Bh = 2s tan 1 3 where 2b is the residual backlash of the NTCIG with backlash control and adjustments to the axial displacement of the driven gear. Based on the definition of the DTE, we have, Therefore, the factors for the backlash, drive/backside mesh stiffness, external load, and axial movement all affect the DTE of the NTCIG based on Eqs.\u00a0(14)\u2013(19) which can be solved using the Runge\u2013Kutta method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000003_iros40897.2019.8968039-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000003_iros40897.2019.8968039-Figure6-1.png", + "caption": "Fig. 6. The experiment setup. (a) Tele-operation system. (b) Various configurations of sheath used in the experiment. 1) 0\u00b0. 2) 90\u00b0. 3) 180\u00b0. 4) 270\u00b0.", + "texts": [ + " The rendered images in \ud835\udd4f\ud835\udd4f\ud835\udc52\ud835\udc52 labeled with the actual joint angles \ud835\udf03\ud835\udf031 and \ud835\udf03\ud835\udf032 of the surgical instrument are pre-transformed into feature vector \ud835\udc53\ud835\udc53x2 using the learned network model. The equation used in searching for the best candidate is given by the following, Eq. (17). \ud835\udc36\ud835\udc36(\ud835\udc65\ud835\udc651,\ud835\udd4f\ud835\udd4f\ud835\udc52\ud835\udc52) = \ud835\udc5a\ud835\udc5a\ud835\udc52\ud835\udc52\ud835\udc4e\ud835\udc4e\ud835\udc5a\ud835\udc5a\ud835\udc51\ud835\udc51\ud835\udc51\ud835\udc51\ud835\udc50\ud835\udc50\ud835\udc51\ud835\udc51(\ud835\udc65\ud835\udc651 \u2218 \ud835\udc65\ud835\udc652)2, \ud835\udc65\ud835\udc652 \u2208 \ud835\udd4f\ud835\udd4f\ud835\udc52\ud835\udc52 , \ud835\udc65\ud835\udc651 \u2208 \ud835\udd4f\ud835\udd4f\ud835\udc4f\ud835\udc4f (17) where \ud835\udc51\ud835\udc51(\ud835\udc65\ud835\udc651 \u2218 \ud835\udc65\ud835\udc652)2represents same as equation (14). Finally, the estimated angles \ud835\udf03\ud835\udf03\ufffd1and \ud835\udf03\ud835\udf03\ufffd2 are determined as the label of the selected image \ud835\udc65\ud835\udc652. The testbed of the flexible endoscopic surgery robot system and master device for tele-operation were used for evaluation as shown in Fig. 6(a). Experiments were carried out in the range of \ud835\udf03\ud835\udf031and \ud835\udf03\ud835\udf032 as defined in a training dataset. The actual joint angle of the surgical instrument is measured by an encoder attached to each joint for reliable measurements. Three different printed images of organ {Colon, Stomach, Larynx} were used as the background. A 2-DOFs flexible surgical robot testbed driven by a tendon-sheath mechanism using a SUS spring and stainless steel wire was used to evaluate the performance of the proposed method. The main driving part consisted of four motors (Dynamixel xm430, ROBOTIS, Korea) and was connected via ball screws and linear guides together", + " The proposed pose estimation method may have some limitations in real surgical environment where optical illumination, dynamic background and surgical smoke can affect the image segmentation outcome. The effect of these factors in pose estimation will be further evaluated. Moreover, a more robust image segmentation method using deep learning will be tried in our further works. To validate the proposed compensator to reduce the hysteresis in various sheath configurations, four different sheath configurations were used as shown in Fig. 6(b). Seven trials periodic sinusoidal signals were given. As a measurement of evaluation, the hysteresis size defined as the peak-to-peak error between the desired trajectory and measured output was used. In addition, the RMS deviation between the desired angle and the actual angle measured by an encoder attached at the joint was also utilized. The hyperparameters \ud835\udc3e\ud835\udc3e\ud835\udc5d\ud835\udc5d, \ud835\udc3e\ud835\udc3e\ud835\udc56\ud835\udc56, and \ud835\udc3e\ud835\udc3e\ud835\udc51\ud835\udc51 were determined by manual tuning. The results are summarized in table \u2161 and illustrated in Fig. 8. In all four cases, the size of hysteresis was reduced to less than 10\u00b0 by averaging the results at \ud835\udf03\ud835\udf031 and \ud835\udf03\ud835\udf032 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001110_6.2014-4161-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001110_6.2014-4161-Figure1-1.png", + "caption": "Figure 1. Relative frame and curvilinear states. The z\u0302 direction completes the right-hand frame.", + "texts": [ + " The y\u0302 axis completes the right-hand frame. In the following of the paper, in-plane mean relative curvilinear coordinates, w\u0303 = (x\u0303, y\u0303, v\u0303x, v\u0303y) T , are exploited instead of the Cartesian mean relative position and velocity, w = (x, y, x\u0307, y\u0307) T : x\u0303 = \u221a (rt + x) 2 + y2 \u2212 rt v\u0303x = x\u0307 cos\u03b1\u2212 y\u0307 sin\u03b1 y\u0303 = \u221a (rt + x) 2 + y2 \u03b1 v\u0303y = x\u0307 sin\u03b1+ y\u0307 cos\u03b1 (1) where rt is the current mean radius of the target's orbit, and \u03b1 = tan\u22121 y rt+x . The relative frame and the graphical interpretation of the curvilinear coordinates are illustrated in Figure 1. The rendez-vous maneuver consist in the solution of the following two-point boundary value problem: dw\u0303 dt = \u2207w x\u0307 (w\u0303) y\u0307 (w\u0303) f (w\u0303, t, \u03b4) \u00b7 x\u0302 f (w\u0303, t, \u03b4) \u00b7 y\u0302 + 0 0 0 1 \u2206fd (w\u0303, t, \u03b4, u) , w\u0303(0) = w\u03030, w\u0303 (tf ) = 0; (2) here: \u2022 w\u03030 and tf are the initial conditions and the maneuvering time, respectively; \u2022 f (w\u0303, t, \u03b4) consists of all the non-inertial accelerations and di erential forces per unit mass but the di erential drag contribution, which is expressed by the term \u2206fd de ned as: \u2206fd = fd,c \u00b7 (sin\u03b1x\u0302 + cos\u03b1y\u0302)\u2212 fd,t \u00b7 y\u0302 (3) 1I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002943_s00202-020-00957-0-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002943_s00202-020-00957-0-Figure8-1.png", + "caption": "Fig. 8 Magnetic flux distribution for a SCIM and b WRIM, with 50% rotor eccentricity", + "texts": [ + " As the additional magnetic flux waves caused by rotor eccentricity are not originated from the same harmonic group of the fundamental pole pair harmonic, the series-connected rotor winding in a WRIM prevents the flow of circulating current or additional rotor current which is induced by the additional flux harmonics. On the contrary, additional magnetic flux waves can induce additional rotor current in the parallel-connected rotor bar. As the induced current in SCIM will produce counteracting flux that damps the additional magnetic flux wave, the magnetic flux wave of a SCIM with an eccentric rotor will remain almost the same as a healthy machine. The comparison between SCIM and WRIM is shown in Fig.\u00a08, where this UMP damping effect is extensively discussed in [23]. The iron losses mainly include hysteresis and eddy current loss. Hysteresis loss is caused by the energy dissipation from the domain changes of the ferromagnetic material, whereas the magnetic field experience changes in the ferromagnetic material. However, the nonhomogeneous flux distribution in the material may cause additional losses that could not be fitted in the hysteresis loss curve, so the anomalous loss is introduced to explain this scenario" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003523_j.triboint.2020.106664-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003523_j.triboint.2020.106664-Figure1-1.png", + "caption": "Fig. 1. Codex Atlanticus, f. 17v, Biblioteca Ambrosiana, Milan. Copyright Biblioteca Ambrosiana.", + "texts": [ + " Based on geometric tools and his talent in the strokes, he determined quite precise shapes and dimensions for lenses and mirrors. To obtain concave surfaces, Leonardo designed various types of machines as reported in Codex Atlanticus and Codex Madrid I. He proposed that machines for polishing mirrors can be classified as machines for casting moulding and machines for metal sheet forming. The machines for casting moulding are found in the Codex Atlanticus, folio 17v, 879r and 1057v. Specifically, the machine presented in folio 17v, see Fig. 1., shows some tribological aspects to analyse. According to the report of Dupre [12] the machine was supposed to make, \u201cparts of balls which were filled with burning materials and thrown or shot from cannons, it is much more likely that the \u2018spheres\u2019 were indeed mirrors\u201d. The machine movement can be described as: \u201cA horizontally mounted crank wheel turns the concave sphere (S in Leonardo\u2019s drawing) which is pressed against a grinding stone T (which in the annotation to one drawing is said to be made of iron)\u201d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002235_s11661-016-3658-5-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002235_s11661-016-3658-5-Figure8-1.png", + "caption": "Fig. 8\u2014Schematic of wrap bend testing and pure bending setup, (a) longitudinal section of the sheet under wrap bend test; (b) and (c) pure bending setup in the model.", + "texts": [ + " The single-crystal plasticity model, along with the failure behavior is METALLURGICAL AND MATERIALS TRANSACTIONS A implemented into the finite element software ABAQUS[11] through a user materials subroutine UMAT. Rather than directly modeling the wrap bend test, the pure bending computational scheme is selected in order to avoid introducing friction into the model, so that attention can be focused on the AA5754-O sheet itself. We consider a sheet with an initial thickness h0 and length 2L0 subjected to pure bending by rotating its edges through an angle 2h as shown in Figure 8. The top \u00f0r0\u00de and bottom \u00f0ri\u00de surfaces of the sheet are required to remain traction free so that Tx \u00bc Ty \u00bc 0 at y \u00bc 0 Tx \u00bc Ty \u00bc 0 at y \u00bc h0 ; \u00bd6 where Tx; Ty are surface traction in x and y directions. The edge of the sheet x \u00bc L0 is rotated through an angle h relative to the center line but still remains shear METALLURGICAL AND MATERIALS TRANSACTIONS A free. Taking the point about which the linex \u00bc L0 rotates to be y \u00bc h0 2 , the bending angle in terms of the current Cartesian coordinates can be written as x L0 y 1 2 h0 \u00bc tan h \u00bd7 Followed Triantafyllidis and Needleman\u2019s approach,[12] Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001360_auto-2013-1038-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001360_auto-2013-1038-Figure3-1.png", + "caption": "Figure 3: Lateral dynamics of a test airplane.", + "texts": [ + " Synchronization of test airplanes (Grumman X-29) in lateral dynamics is considered, which is a typical benchmark example for the synchronization problem. The same synchronization problem has been also considered in [7]. The linearized plane model is introduced in [1] and is given by equation (1) with the matrices \ud835\udc34 =( \u22122.59 0.997 \u221216.55 0 \u22120.1023 \u22120.0673 6.779 0 \u22120.0603 \u22120.9928 \u22120.1645 0.04413 1 0.07168 0 0 ) , \ud835\udc35 =( 1.347 0.2365 0.09194 \u22120.07056 0.0006141 0.0006866 0 0 ) and with the subsystem state \ud835\udc65 \ud835\udc56 (\ud835\udc61) = (\ud835\udc5d \ud835\udc56 , \ud835\udc5f \ud835\udc56 , \ud835\udefd \ud835\udc56 , \ud835\udf19 \ud835\udc56 ) . As shown in Figure 3, \ud835\udc5d describes the roll rate of the plane, \ud835\udc5f the yaw rate, \ud835\udefd the roll angle and \ud835\udf19 the yaw angle. Clearly, a complete synchronization in all four state variables will be achieved. In the further pictures, synchronous behavior only in the first state variable will be shown, since the behavior in the other states are also similar. It is assumed that the planes are located on a string and an information exchange between the nearest planes is considered. Consequently the Laplacian matrix has the following structure: \ud835\udc3f = ( 1 \u22121 0 0 \u22121 2 \u22121 0 0 \u22121 2 \u22121 0 0 \u22121 1 ) The continuous-time controller\ud835\udc3e os = \ud835\udc3f \u2297 \ud835\udc3e with the controller matrix \ud835\udc3e = ( 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000927_2014-01-0880-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000927_2014-01-0880-Figure1-1.png", + "caption": "Figure 1. Virtual vehicle frame with position of regarded mounts and bushings marked.", + "texts": [ + " In the standard development process, amplitude dependence due to the Payne effect is not considered within the MBS environment but simple spring-damper models are used for cab mounts, suspension bushings and jounce bumpers. Only the non-linear static stiffness curves are implemented in standard MBS models. For this research, new non-linear bushing models were used for all cab mounts within the ADAMS/Car environment to assess the potential of improvement for full-vehicle simulation of durability loads. Three different types of mounts are regarded, Figure 1 showing their position in the car frame. Overall, six mounts are modeled with the new non-linear model approach, with different characteristics in all three translational directions. For parameter identification the mounts have been tested under quasi-static and sinusoidal excitation. The models were then validated under realistic transient excitation derived from on-road experiments. Simulation accuracy and effort for parameter identification will be evaluated based on these experiments. The term elastomer refers to any rubber parts including synthetic rubber compounds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003586_s12206-020-0918-5-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003586_s12206-020-0918-5-Figure7-1.png", + "caption": "Fig. 7. Distribution of the minimum velocity evaluation index.", + "texts": [ + " And the output maximum velocity evaluation index of the 3 - PPRU PM with an inconstant constant Jacobian grows at an exponential rate. Figs. 5 and 6, respectively, show the contour of the maximum velocity evaluation index of the PPRU 2PPRU+ PM and 3 - PPRU PM. The maximum velocity evaluation index of the PPRU 2PPRU+ PM increases slightly along Py-axis and keeps constant along Px-axis. While the maximum velocity evaluation index of the 3-PPRU PM increases smoothly in a certain workspace and begins to rocket rapidly around Px = \u00b1 0.5 m and Py = \u00b1 0.5 m nearby. Fig. 7 describes the minimum velocity evaluation index of the PM with a completely constant, partially constant and inconstant Jacobian. The PM with a completely constant Jacobian can always maintain a constant minimum velocity evaluation index about 0.9428. The output minimum velocity evaluation index of the PM with a partially constant Jacobian decreases smoothly. The output minimum velocity evaluation index of the PM with an inconstant constant Jacobian has a moderate downtrend. More specifically, from Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000482_acc.2015.7171930-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000482_acc.2015.7171930-Figure1-1.png", + "caption": "Fig. 1. The FSD ExtremeStar", + "texts": [ + " Due to the addition of the feedback term in the reference model, the CRM trajectory deviates from its ORM counterpart and towards the plant output, which reduces the error between the reference model and the plant. Thus, the burden on the adaptation is reduced, which can lead to less oscillatory adaptation. In the following application it is shown that the CRM adaptive system also provides additional robustness to unmodeled dynamics. The presented control architecture is applied to the FSD ExtremeStar, a modified version of the model airplane Multiplex TwinStar II, which is shown in Fig. 1. The modified aircraft possesses 16 different controls in total. The presented application only utilizes the three conventional controls: asymmetrical aileron deflection \u03be, symmetrical elevator deflection \u03b7, and rudder deflection \u03b6. A nonlinear six degree of freedom simulation model was developed in [18] and is used in this study. The simulation model provides a highly realistic environment to develop and test the controller. It includes modeling of the nonlinear equations of motion, nonlinear aerodynamics with propeller wing coupling, an environment model based on the international standard atmosphere, as well as actuator and sensor models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000171_ijmmm.2019.098066-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000171_ijmmm.2019.098066-Figure2-1.png", + "caption": "Figure 2 (a) CNC turning of Inconel 825 (b) MQL machining setup (see online version for colours)", + "texts": [ + " The one end of Inconel 825 bar was held in the machine chuck and other end was supported by tailstock so as to minimise run out and vibration during machining experiments. The length of machining used for experimental runs was 100 mm. During MQL machining, commercially available lubricant (LRT 30) was supplied through a nozzle at a flow rate of 60 ml/hr with 5 bar air pressure. The MQL delivery system was placed at tool rake face and the nozzle was kept at 2 mm away from the tool tip (Ali et al., 2015). Figure 2(a) shows the experimental set up used for turning of Inconel 825 in this research and Figure 2(b) shows the MQL machining setup employed in the present work. Three variables with three levels comprising 15 experimental runs following a Box-Behnken design shown in Table 2 are employed. The surface finish of the specimen was measured with stylus-based contact type profilometer (Make: Taylor Hobson, Surtronic 2.5, UK) having a diamond stylus of 5 \u00b5m tip radius and with cut off length of 0.25 mm. Three measurements were taken during each experimental run. The values are found very close to each other and the average was considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001502_s0263574714002276-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001502_s0263574714002276-Figure2-1.png", + "caption": "Fig. 2. (Colour online) Schematic view of a single actuator.", + "texts": [ + " Moreover, the generalized coordinate vector describing the position and orientation of the movable platform is denoted \u03c7= [x, y, z, \u03c6, \u03b8, \u03c8]T (which is the Euler angle description). The following subsections briefly describe the kinematics, multi-body dynamics, permanent magnet synchronous motor (PMSM) system dynamics, and the control problem statement for the uncertain electrical Stewart platform. http://journals.cambridge.org Downloaded: 11 Mar 2015 IP address: 130.126.162.126 Adaptive vector sliding mode fault-tolerant control of Stewart platform 3 Each motor drives an actuator (illustrated in Fig. 2) by a synchronous belt and screw rod. In Fig. 2, ac and bc are the mass centers of the moving actuator part and the rotating actuator part respectively; va,i , vac,i , and vbc,i are the velocities of the upper joint, ac point, and bc point respectively; \u03c9a,i and \u03c9m,i are the angular velocities of the upper joint and rotor respectively; rac and rbcare the distance from ac to the upper joint and the distance from bc to the lower joint respectively; and i refers to the ith actuator. Kinematic analyses of the Stewart platform can be divided into analyses of the inverse kinematics and forward kinematics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000919_vppc.2014.7007108-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000919_vppc.2014.7007108-Figure2-1.png", + "caption": "Fig. 2. A clutch disc [4].", + "texts": [ + " The diaphragm spring is responsible for the clamp load and the pressure plate for applying this force on the disc. Considering a good level of friction between the disc and the pressure plate, the torque can be transmitted from the motor to the gearbox. 978-1-4799-6782-7/14/$31.00 \u00a92014 IEEE A. Some Details of the Facing Facing is the part of the disc responsible for the transmission of the torque from the engine to the gearbox by the coefficient of friction and is usually composed of thermoset resins (e.g. phenolic), rubber, fillers and fibers like glass, metallic and polymeric. Figure 2 shows a disc of a dry clutch where the facing is normally fixed by rivets, as illustrated in Figure 3. In general, the facing is produced from 160 mm to 430 mm outside diameter, depending on the vehicle size. The most important characteristic of the facings is the reliable level of friction in different work conditions of the vehicle. B. The Judder Behavior Clutch judder [5] is the term used to describe longitudinal oscillation during the clutch slip phase in vehicles when accelerating. The frequency of such automotive behavior ranges from 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003308_s40430-020-02488-y-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003308_s40430-020-02488-y-Figure2-1.png", + "caption": "Fig. 2 The tooth flank formation mechanism of LSBG", + "texts": [ + " Then, the sweep line (logarithmic spiral line) of the gear is set up. Finally, the scanning section is scanned along the sweep line to get the gear. A three-dimensional model of LSBG can be obtained through a tooth array. The tooth profile surface of LSBG can be regarded as a spatial curved surface, which formed by the spatial motion of the logarithmic spiral line on the plane, when the plane (meshing surface) is purely rolling along the base cone. The tooth flank formation mechanism of LSBG is shown in Fig.\u00a02. Recently, Shang [5] used the meshing principle and transmission characteristics of logarithmic spiral bevel gears to derive the logarithmic spiral equation. Afterward, the involute equation was also derived by using the principle Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:400 1 3 Page 3 of 14 400 of involute formation. The conic logarithmic spiral line is shown in Fig.\u00a03. In the Pro/E environment, these two curves were set up, and the logarithmic spiral line was projected on the degree circle, and they were later scanned and mixed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002246_978-3-319-15684-2-Figure14.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002246_978-3-319-15684-2-Figure14.3-1.png", + "caption": "Fig. 14.3 Critical values of the parameter of dissipation", + "texts": [ + " At the constant value of the first transfer function P0 \u00bc const, we have P00 \u00bc 0 and d 1 [ 0 that, as expected, corresponds to damped free vibrations during the entire kinematic cycle. To estimate the energy variation intensity let us make use of the function E \u00bc p2\u00f0t\u00deS2\u00f0t\u00de; \u00f014:12\u00de to which the amplitude values of the energy are proportional. Based on Eqs. (14.8) and (14.12), the condition dE/dt < 0 yields the following result: n _p=\u00f02p\u00de[ 0: \u00f014:13\u00de According to Eqs. (14.5) and (14.13), the condition which is analogous to relation (14.11) takes on the form d[ d \u00bc 3l2P0 P 00 x0 2k1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe l2P02 q : \u00f014:14\u00de Figure 14.3 shows the family of critical values of the d \u00f0u0 ; l 2\u00de parameter. The energy of vibratory systems increases in the process of acceleration (P0 P 00 > 0) and decreases during deceleration \u00f0P0 P 00 \\0\u00de. As \u03bc approaches unity, the intensity of vibrational energy variations decreases. Let us now return to the analysis of the initial model with two degrees of freedom (Eq. 14.2). As a rule, the parameter n in problems of the dynamics of mechanisms negligibly affects the \u201cnatural\u201d frequency p and, at the same time, 122 I", + " The analysis showed that this is related to formation of the beatmodewhen thework of gyroscopic forces vanishes only after the beat period (which often exceeds the period of the kinematic cycle s \u00bc 2p=x0 by one order of magnitude) is over. Figure 14.6c plots the E0 \u00bc dE=du0 function which characterizes the intensity of energy variations for mode 4 (k1 \u00bc 10; k2 \u00bc 30; l2 \u00bc J2=J1 \u00bc 0:2). This mode is interesting since, in fact, the situation considered above for the model with one degree of freedom (see Fig. 14.3) which corresponds to the limiting case when k22=k 2 1 1 is repeated. The comparison of the E0 plot with that shown in Fig. 14.3 shows that the plot of d \u00f0u0; l\u00de obtained at the limiting transition serves as an analogue for the envelope for high frequency vibrations that were identified when the elasticity of the output unit was taken into account. This study shows that the energy accumulated in the vibratory system in cyclic machines is mainly determined by the work performed by the external source upon implementation of nonstationary 14 Vibrations Excitation in Cyclic Mechanisms \u2026 125 links that manifests itself most clearly in local violations of stability conditions at certain sections of the kinematic cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002915_s10846-020-01168-2-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002915_s10846-020-01168-2-Figure6-1.png", + "caption": "Fig. 6 The configuration of a robot walking helper", + "texts": [ + " proposed learning-based control schemes for humanoid robots [19, 20]. And, Ko et al. proposed a braking controller using differential flatness for the passive type of robot waking helper [16]. The proposed control strategy was developed basically following that in [16], with modification for adapting to the active one. Before the derivation of the control strategy for executing the planned path, we first briefly introduce the dynamic model of an active type of robot walking helper, based on the configuration shown in Fig. 6. In Figure 6, fh and \u03c4h, as the inputs, are the force and torque exerted by the user, respectively, \u03c4r and \u03c4l, as the outputs, are the torque to be generated by the right and left motors, v and \u03c9 the linear and angular velocities of the robot walking helper, and \u03b8r\u02d9 and \u03b8l\u02d9 the angular velocities of the right and left wheels. In Cartesian coordinates, its state is described as q \u00bc x; y; \u03b8\u00bd T \u00f06\u00de where [x, y, \u03b8]T stands for the position and orientation. Under the assumption of non-slip at the wheel contact point on ground, the nonholonomic constraints are given as" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003784_j.ejps.2020.105649-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003784_j.ejps.2020.105649-Figure8-1.png", + "caption": "Fig. 8. Concentrations of simvastatin (SIM) and simvastatin hydroxyacid (SIM HA) formed from: simvastatin powder, drug-loaded PVP particles and drug-loaded PVP/BHA particles in citrate (CB) and phosphate buffers (PB). The samples were analysed after 2, 24 and 48 hours of mixing.", + "texts": [ + " The same but more pronounced effect was observed for drug-loaded PVP particles in PB, where after 24 hours of mixing, the concentration of simvastatin in PB decreased more radically than in CB buffer. This possibly happens because of the pH difference between the buffers and can be related to the simvastatin hydrolysis moiety - active metabolite, SIM HA, where the observed trend in SIM HA concentrations is opposite to that of simvastatin. To be exact, we evaluated the concentration of SIM HA during solubility studies and measured the change in SIM HA concentration with time in both buffers and observed its relation to simvastatin concentration (Fig. 8). Incorporation of simvastatin into the drug-loaded PVP particles increased SIM HA formation in comparison to pure simvastatin. B. Sterle Zorec et al. European Journal of Pharmaceutical Sciences 158 (2021) 105649 Moreover, the SIM HA concentration is increasing with time, while the simvastatin concentration is decreasing. According to our results the higher the pH, the higher the SIM HA concentration, which is opposite to the outcome we observed with simvastatin concentrations as described previously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000351_icuas.2015.7152431-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000351_icuas.2015.7152431-Figure2-1.png", + "caption": "Figure 2. Sheme of Quadrirotor", + "texts": [ + " This UAV with a rotary wing has four motor arranged in the form of a cross of equal sides. Its structure offers him the advantage to take off and land vertically, and also to carry out stationary flights. The roll and pitch moments are obtained by varying the rotational speeds of two opposite motors, furthermore, two of the four rotors have a reverse pitch propeller which allows to cancel the yaw moment. The different moments and forces due to the engines taken into account in this study are shown in Figure 2. Thus, each of the four engines \ud835\udc40\ud835\udc56 produces \ud835\udc39\ud835\udc56 force and \ud835\udf0f\ud835\udc56 torque on the\ud835\udc67 axis. The total thrust is the sum of the forces of the four actuators \ud835\udc62 = \ud835\udc391 + \ud835\udc392 + \ud835\udc393 + \ud835\udc394. The torque \ud835\udf0f\ud835\udc65around the x axis is obtained by the difference of the forces and \ud835\udc392 \u2212 \ud835\udc394and the torque \ud835\udf0f\ud835\udc66 around the y axis by the difference \ud835\udc391 \u2212 \ud835\udc393forces. Finally, the torque \ud835\udf0f\ud835\udc67 around the \ud835\udc67 axis is obtained by the sum of the torques produced by the motors \ud835\udf0f1 + \ud835\udf0f3 \u2212 \ud835\udf0f2 \u2212 \ud835\udf0f4, \ud835\udc401 and \ud835\udc403turning effect in the positive direction while \ud835\udc402 et \ud835\udc404 rotate in opposite directions (see Figure 2). The dynamic model of Quadrirotor is calculated using the Euler-Lagrange formalism [18], [19] and can be presented in state space form of : ?\u0307? = \ud835\udc53(\ud835\udc4b, \ud835\udc48) by introducing as state vector: \ud835\udc4b\ud835\udc47 = [\ud835\udf11, \ud835\udf11,\u0307 \ud835\udf03, \ud835\udf03,\u0307 \ud835\udf13, \ud835\udf13,\u0307 \ud835\udc65, \ud835\udc65,\u0307 \ud835\udc66, ?\u0307?, \ud835\udc67, \ud835\udc67,\u0307 ]. The control of Quadrirotors is achieved through a conventional sliding mode control, the latter noted \ud835\udc481, \ud835\udc482, \ud835\udc483, \ud835\udc484, \ud835\udc48\ud835\udc65 and \ud835\udc48\ud835\udc66 are supposed to be already accomplished [18]. In order to perform a flight training fleet, each Quadrirotor has calculate its desired trajectory at every step time, as illustrated in Figure 3, such that : \ud835\udc65\ud835\udc56\ud835\udc51 (\ud835\udc61 + 1) = \ud835\udc65\ud835\udc56 (\ud835\udc61) + \u210e , where \u210e \u2208 \ud835\udc453 is the vector that minimizes best the equation (10)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000623_s00542-015-2422-x-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000623_s00542-015-2422-x-Figure1-1.png", + "caption": "Fig. 1 Principle of the SGCMG driven by hollow USM", + "texts": [ + " It can achieve high-precision servo speed control in aerospace systems and can improve the systems effective load for it\u2019s simple structure, fast response, high angular resolution, low speed with high speed, self-locking, and without magnetic interference. In this paper, a control system of SGCMG driven by hollow USM is designed to achieve high stability speed control. To meet the requirements of the control indicators of the system the robust control theory based on fuzzy T-S theory was introduced. The whole control system will realize on actual SGCMG driven by a hollow USM on the basis of simulation. The working process of SGCMG driven by a hollow USM is shown in Fig. 1, where m is the installation coordinates of SGCMG, and f is the fixed coordinates of SGCMG. The momentum wheel mounted on orthogonal the gimbal frame, and rotating around S-axial with the driving of rotor. The frame g-axis is perpendicular to the momentum t-axis. The g-axis with the s-axis and the t-axis constitute the right-handed orthogonal coordinate system. Single-gimbal control moment gyro (SGCMG) is an attitude control device by a spinning rotor and one motorized gimbals that tilt the rotor\u2019s angular momentum (Masterson et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000584_j.jsc.2014.09.031-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000584_j.jsc.2014.09.031-Figure13-1.png", + "caption": "Fig. 13. Knot of dodekagon-like shape.", + "texts": [ + " A method of Gr\u00f6bner bases computation of polynomials with approximate numeric (floating point) coefficients such as developed by Sasaki (2012) will be of great help to our research. We see several directions of the research based on our results and methodologies. We are aware that Brunton (1961) gave several interesting examples without proofs. With the use of Eos it will become a routine work to generate polygonal knots for these examples. Although theoretically not significant, the polygonal knots of even numbered edges (cf. Fig. 13) are interesting to observe. It is generated by the method of the algebraic constraint solving discussed in Section 4.3. Maekawa (2011) gave visionary introductory accounts on knot folds. We believe that to establish origamists\u2019 ideas firmly on the ground of computer science and mathematics would be an important step to direct. Alexander, J.W., Briggs, G.B., 1926. On types of knotted curves. Ann. Math. 28 (1/4), 562\u2013586. Alperin, R.C., Lang, R.J., 2009. One-, two, and multi-fold origami axioms. In: Origami4, Proceedings of the Fourth International Meeting of Origami Science, Mathematics, and Education" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003855_ddcls49620.2020.9275206-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003855_ddcls49620.2020.9275206-Figure2-1.png", + "caption": "Fig. 2: Sinusoidal wave pattern", + "texts": [ + " It possess two biomimetic propellers to achieve multimodal motions, a 5-DoF (degree of freedom) manipulator for underwater operation and two pairs of binocular cameras for underwater detection and observation. Some actual measured parameters of the UBVMS are listed in the Table 1. The biominetic underwater propeller possesses a flexible undulating fin which consists of twelve short rays connected by a black silicone sheet. By distributing all rays in interval phases of a sinusoid, propulsive force can be generated by the propeller. As shown in Fig. 2, the motion model of a single fin ray is: \u03b8(s, t) = \u0398 sin 2\u03c0( s \u03bb + ft+ \u03c6) (1) where \u03b8(s, t) defines the angular deflection of the ray at distance s along the axis X of coordinate system O-XY Z at time t. \u0398 is the maximum angular deflection of the sinusoidal waves. \u03bb is the wavelength. f is the frequency of the waves. 2\u03c0\u03c6 is the initial phase of the waves. When value of f is not 0, with the time t goes, a wave is considered propagating along the axis X . The force generated by this wave can be divided into two types: thrust T along the axis X and lateral force L along the axis Z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001917_ccdc.2014.6852195-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001917_ccdc.2014.6852195-Figure3-1.png", + "caption": "Fig 3. Four aerodynamic configurations", + "texts": [ + " Suppose the uncertainty coefficients, corresponding to 1\u03b8 , 1\u03b8 , 1\u0394 , 1\u0394 ,CL, eLC \u03b4 , CD, Cm, , em mC C\u03b4 \u03b1 , and mqC , follow normal distribution with mean 1 and standard deviation 0.2. The uncertain parameters used in SRAD is calculated by 0 i i ia v a= (16) where 0 ia is the standard value of the i-th uncertain parameter. 2) The control law design For the morphing aircraft discussed in this paper, the controller\u2019s design vector d is designed based on SRAD for two boundary aerodynamic configurations ((a) and (b) as shown in Fig.3). The controller\u2019s two sets of parameters [kpcK kp ki] are worked out. The parameters used in other configurations are got using Matlab\u2019s interpolation function newrb based on their span and sweep. In the design process, V is supposed to be constant, which can be got by control the engine. The system\u2019s state vector x= [ q h e], the control input u = c, the command input is *. The control law is * * 0 ( ) ( ) t c pc p ik k k dt\u03b4 \u03b8 \u03b8 \u03b8 \u03b8= \u2212 + \u2212 + \u2212Kx (17) 5 CONTROL SIMULATION 5.1 Simulation Platform and Parameters This paper chooses the Teledyne Ryan BQM-34 \u201cFirebee\u201d as the simulation base platform", + " The main simulation parameters are shown in Table 1. Based on the quasi-steady assumption, the software Missile Datcom[23] is employed to calculate the morphing aircraft\u2019s aerodynamic coefficients. NACA-0010 standard airfoil is adopted. The taper ratios of the wing\u2019s inner part and outer part are 1. When the wing is not swept, the length of chord the two parts is 0.8m and 0.7m. 476 2014 26th Chinese Control and Decision Conference (CCDC) Aerodynamic coefficients of four different configurations are calculated (as shown in Fig 3). Configuration (a): the wings are not swept and the outer parts are fully stretched. Configuration (b): the wings are swept 15 and the outer parts are 2/3 stretched. Configuration (c): the wings are swept 30 and the outer parts are 1/3 stretched. Configuration (d): the wings are swept 45 and the outer parts are fully folded. Linear interpolation between these four different configurations was used to construct the aerodynamic data during transition from configuration (a) to (d). 5.3 SRAD Design Results The two different configurations, (a) and (d) in Fig. 3, are suitable for loiter and dash respectively. The design point is the aircraft\u2019s steady-state flight condition when Ma = 0.5, h= 10000m, and the aircraft is trimmed at =2.3 and =4.3 respectively. Corresponding to control law (17), the design results for the control law\u2019s parameters are: For aerodynamic configuration (a): [ 47.92 9.377 83.73 0.5943 0.3533] 0.932 0.1863 pc p i k k k = \u2212 \u2212 \u2212 \u2212 = =\u2212 K For aerodynamic configuration (d): [747.8 30.99 770.7 4.711 2.895] 0.1534 0.3192 pc p i k k k = \u2212 \u2212 \u2212 = \u2212 =\u2212 K 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003974_s12239-021-0023-5-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003974_s12239-021-0023-5-Figure2-1.png", + "caption": "Figure 2. Pressure distribution test in contact area.", + "texts": [ + " Pressure Distribution Measurement in Contact Area Tekscan\u00ae pressure measurement system was used to test the pressure distribution in the contact area and as well obtain the contact patch of test tire, which was then compared with the contact patch obtained by the subsequent deformation distribution test to verify the effectiveness of the screening method of the contact area. Table 2 lists the parameter configuration of Tekscan pressure measurement system used in this paper. During the test, the inflation pressure of all tires was maintained at a rated value 250 kPa, and the rated vertical load was 500 kg. Tire loading was performed on the stiffness test table (as shown in Figure 2) using the lifting and loading mechanism. 2.3. Deformation Distribution Measurement in Contact Area The VIC-3D non-contact strain measurement system based on digital image correlation technology (as shown in Figure 3) was used to obtain the three-dimensional coordinates of the tread in the contact area after loading, as well as the radial, transverse and longitudinal displacement with accompanying strain. Table 3 lists the parameter configuration of VIC-3D measurement system. The test conditions were consistent with the pressure distribution test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001394_mtsj.47.3.4-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001394_mtsj.47.3.4-Figure2-1.png", + "caption": "FIGURE 2", + "texts": [ + " f \u00bc l d ; AR \u00bc b2 S ; \u03ba \u00bc b d ; and \u223ct \u00bc t c We also consider a few other configuration parameters for the wing and vertical stabilizer, including wing longitudinal position lw, vertical stabilizer longitudinal position lv and area of the vertical stabilizer Sv. The aerodynamic center of the wing and the vertical stabilizer, which is located near the quarter-chord line, are \u223c lw and \u223c lv (normalized by the fuselage length l ) aft of the center of buoyancy, respectively, while the area of the vertical stabilizer Sv is normalized by the fuselage frontal area Sf . All of the geometric parameters that we consider are indicated in Figure 2. A critical parameter for performance analysis is the buoyant lung capacity \u03b7, given as a percentage of the neutrally buoyant displacement. Buoyancy actuators include pistoncylinder arrangements, oil-filled bladders, and pneumatic bladders. For simplicity, we consider the case of a piston-cylinder actuator and assume a cylindrical hull with constant crosssectional area (see Figure 3). Recall that the dynamic model presented earlier assumes changes in buoyant volume occur at the origin of the body frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003172_14484846.2020.1769462-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003172_14484846.2020.1769462-Figure4-1.png", + "caption": "Figure 4. The assembled solid model and meshed model of harmonic drive without the ring gear.", + "texts": [ + " First, the strains on FG cup surfaces are recorded due to split-cam SWG insertion with various split cam of different geometries, and secondly, strains are recorded on FG cup with fully assembled HD and applied load. The strains are then compared with FE simulation. For complete stress mapping on the flex gear, FE analysis is carried out. First the flex-gear and split cam SWG assembly is modelled. There are cup-shaped flex gear, ball bearing with flexible races and split cam in the assembly. The static structural analysis in FE is carried out. The assembly model is meshed for split cam SWG insertion only as well as full applied loading with cam insertion as shown in Figure 4. The rigid boss of the flex gear is constrained in all direction motion. Figure 4(b) shows the FE meshed model. The FG cup is meshed with tetrahedron element. However, with hexagonal element, the mesh convergence does not occur as the FG cup is the complex geometry. Simultaneously, the FRB with split cam is meshed with hexagonal element. With the mesh convergence study, the model (Figure 4 (b)) has 253,111 numbers of elements and 130,003 numbers of nodes. All solid bodies are modelled with SOLID186 element with the material properties of structural steel. There are 49 contacting surfaces in the model. These contacting surfaces are between inner surface of Figure 2. The proposed split cam SWG for harmonic drive. Figure 3. The detailed experimental setup for testing harmonic drive with split-cam SWG. the FG cup and outer race of the FRB; the groove in outer race of FRB and the balls; the balls and the groove in inner race of FRB; finally the inner race of FRB and splitcam surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003320_j.cmpb.2020.105646-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003320_j.cmpb.2020.105646-Figure6-1.png", + "caption": "Fig. 6. The articulator in the initial and final poses.", + "texts": [ + "9) When a dental articulator is used to simulate the jaw moveent and analyze the occlusion, it will usually exercise in three ifferent motions: protrusion, left lateral excursion, and right latral excursion. The lateral excursions will involve jaw rotation and e will define the side to which the mandible moves is the workng side and the opposite side from which the mandible moves s the non-working side. Due to the symmetry of the articulator echanism, we only derive the mathematical model for the right ateral excursion. We define the position of the incisal pin tip in he initial and final pose as point P and point Q , as shown in Fig. 6 . T { c T t s E [ [ c l{ { z i g b [ w t C [ a w r p{ { T z i t F \u03b8 F W c o l m ( a a s c a a p R 0 is the transformation matrix from the world coordinate system X 0 , Y 0 , Z 0 } to the right reference coordinate system { X R ,Y R ,Z R } . Beause the axis directions of both coordinate systems are the same, R 0 is a translation from the origin of the world coordinate sysem to that of the right reference coordinate system. The relationhips of [ P ] 0 with[ P ] R and[ Q ] 0 with[ Q ] R are presented as Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001084_s12206-015-0217-8-Figure22-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001084_s12206-015-0217-8-Figure22-1.png", + "caption": "Fig. 22. Three dimensional FE models with different inclination angles of the BMI nozzles: (a) fn = 13.1o; (b) fn = 23.1o; (c) fn = 36.2o.", + "texts": [ + " So, in this study, effects of the geometric variables on the following total stresses were investigated via performing parametric stress analysis: \u00b7The total stress distributions on the inner surface lines of the down-hill side \u00b7The through-thickness total hoop stress distribution from the maximum stress generation point within the weld zone of the down-hill side \u00b7The through-thickness total axial stress distribution from the maximum stress generation point below the weld root of the down-hill side. Fig. 22 represents three dimensional FE models with different inclination angles of the BMI nozzles. Fig. 23 shows variations of the total stresses on the inner surface lines of the down-hill side vs. the inclination angle (tn = 13.38 mm, ro = 19.57 mm, qw = 25o, dw-up = 18.613 mm, dw-down = 31.7 mm, ww = 8.82 mm). Dotted lines mean maximum stress generation locations within the weld zone. As shown in Fig. 23, it is found that maximum values of the axial and hoop stresses increase with increasing the inclination angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000442_coase.2015.7294131-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000442_coase.2015.7294131-Figure4-1.png", + "caption": "Fig. 4: The case of non zero\u2013symmetric admissible angular range. The dashed lines AB and AC delimit the admissible angular range.", + "texts": [ + " (11) Clearly, inequality (11) is of type (10), hence it can be added to the optimization problem while preserving its convexity. The case of non zero-symmetric admissible range, in which q\u2212i \u2264 qi \u2264 q+i , q\u2212i 6=\u2212q+i , (12) may be treated in a similar way. First we denote by \u03b1 the angle such that 2\u03b1 = q+i \u2212q\u2212i . Then we observe that the point O, that lies on the bisector of the admissible angular range, can still be expressed as a linear combination of decision variables, by taking a proper rotation matrix \u2126\u03c6 . Referring to Fig. 4, the point O may be obtained by adding to pi the vector (pi\u2212 pi\u22121) ri+1 ri rotated by an angle \u03c6 . By defining \u2126\u03c6 = [ cos\u03c6 \u2212sin\u03c6 sin\u03c6 cos\u03c6 ] , the constraint for non zero\u2013symmetric angular range may thus be expressed as:\u2225\u2225\u2225\u2225pi+1\u2212 [ pi +\u2126\u03c6 (pi\u2212 pi\u22121) ri+1 ri ]\u2225\u2225\u2225\u2225\u2264 2ri+1 sin \u03b1 2 . (13) Note that \u03c6 depends only on the structure of the robot (i.e. the length of the links and the admissible angular range) and not on the particular configuration. Hence \u2126\u03c6 is constant, rendering the constraints of type (10) and thus preserving the convexity of the problem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure7-1.png", + "caption": "Figure 7. The contact point of curve-face gear pair.", + "texts": [ + " According to the external generating method, the normal vectors and position vectors at meshing point of non-circular gear and curve-face gear should be equal to each other and Figure 5. The relationship of non-circular gear and curve-face gear. at The University of Melbourne Libraries on June 5, 2016pic.sagepub.comDownloaded from can be established as follows rf !\u00f01\u00de \u00f0u1, 1, s\u00de \u00bc rf !\u00f02\u00de \u00f0u2, 2, s\u00de nf !\u00f01\u00de \u00f0u1, 1, s\u00de \u00bc nf !\u00f02\u00de \u00f0u2, 2, s\u00de 8< : \u00f011\u00de Based on the formation of conjugate tooth surface and the space meshing theory, Figure 7 shows the tooth surface 1 is in external meshing with the tooth surface 3, the tooth surface 1 is in external meshing with the tooth surface 2, too. So, the tooth surface 2 and 3 can be in external meshing with each other at the same time. L13 is the contact line of shaper cutter and curve-face gear, and L12 is the contact line of shaper cutter and non-circular gear. At any meshing point P, shown in Figure 7, it is a common tangent point of the three surface instantaneously and keeps pure rolling contact, respectively, the three surface are the surface 1 of shape cutter, the surface 2 of non-circular gear and the surface 3 of curve-face gear. So, it can be guaranteed that the non-circular gear can mesh with the curve-face gear correctly at any moment. Point P is the contact point of the contact line L12 on non-circular gear and the contact line L13 on curve-face gear in the process of transmission, which makes the contact pattern restricted to local and the meshing process more smoothly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001273_ccdc.2013.6561824-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001273_ccdc.2013.6561824-Figure3-1.png", + "caption": "Fig. 3. Tracking control formulation The virtual velocity controller based on the backstepping approach can be defined as:", + "texts": [ + " The design of the cascaded control strategy consists of two parts: an outer loop kinematic controller and 4915978-1-4673-5534-6/13/$31.00 c\u00a92013 IEEE an inner loop dynamic controller. The fault tolerant control problem will be illustrated in Section 3. In this paper, only the horizontal plane motion control is discussed. For UUV the horizontal plane motion control, the desired state of a reference vehicle is described as [ ]T d d d dx y \u03c8=\u03b7 , [ ]T d d d du v r=q . The actual state of UUV is represented by [ ]Tx y \u03c8=\u03b7 , [ ]Tu v r=q . The detailed description can be seen in Fig. 3. [ ]T d x ye e e\u03c8= \u2212 =e \u03b7 \u03b7 is the tracking error in the inertial frame. ( cos sin ) ( cos sin ) ( sin cos ) ( sin cos ) x y d dc c c x y d d c d k e e u e v eu v k e e u e v e r r k e \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 + + \u2212 = = \u2212 + + + + q (2) where ,k k\u03c8 are constant coefficients. As is well known, backstepping method in the kinematic control will cause speed jump problem. Yang developed the first innovative application of bio-inspired neural dynamics model (bio-inspired model) to real-time mobile robot path planning and found it a useful tool to deal with speed jump problem of backstepping method [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003411_1077546320947338-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003411_1077546320947338-Figure1-1.png", + "caption": "Figure 1. Structure of dual mass flywheel (a) schematic diagram of parts of dual mass flywheel, (b) two-stage stiffness arc spring structure, and (c) angular clearance diagram.", + "texts": [ + " Section 1 briefly describes the structure and working principle of the DMF; In Section 2, the nonlinear model of DMF is established; In Section 3, the analytical solution of the dynamic response of DMF is obtained by average method; In Section 4, the nonlinear frequency response of DMF is analyzed by analytical solution; In Section 5, the influence of DMF parameters on nonlinear frequency response is studied; Finally, Section 6 draws some conclusions from the present work. The structure of the DMF is shown in Figure 1(a). It is composed of the first mass, the second mass, the starting ring gear, the transmission flange, the long arc spring, the arc sheath, the friction plate, the sealing disc, and the bear. The starter ring gear is connected to the first mass by an interference fit. The first mass is connected to the flange at the end of the engine crankshaft by bolts, the flange is connected to the second mass by rivets, and the second mass and the clutch assembly are fixedly connected by bolts. The first mass assembly is connected to the second mass assembly by low rigidity arc spring and can rotate relative to each other", + " When the crankshaft of the engine is rotated, the first mass is driven to compress the arc spring through the boss, and the other end of the arc spring pushes the side ears on both sides of the force transfer flange, to drive the second mass to rotate and realize the powertrain of power from the engine to the gearbox (Chen et al., 2019; Zu et al., 2015). The two-stage stiffness arc spring structure is composed of the outer arc spring and the inner arc spring. The outer arc spring has smaller stiffness k1 and the inner arc spring has larger stiffness k2. The structure of the two-stage stiffness arc spring is shown in Figure 1(b). The inner arc spring is nested in the outer arc spring. When the spring is in a free state, the distribution clearance of the inner and outer arc springs is \u03b81. 1. When the engine is operating at idle or under small load, the compression deformation of the outer arc spring is less than \u03b81; thus, the DMF only has an outer arc spring to transmit torque. 2. When the engine is working under heavy load, the compression deformation of the outer arc spring is greater than \u03b81, the inner and outer arc springs of the DMF transfer torque in parallel. The first-stage stiffness and second-stage stiffness are given as follows K \u00bc k1 (1) K \u00bc k1 \u00fe k2 (2) There is an angular clearance \u03b80 in the DMF, as shown in Figure 1(c). The flange does not contact the arc spring, and the torsional stiffness of the DMF is 0. In summary, the nonlinear stiffness model of DMF can be obtained K\u00f0\u03b8\u00de \u00bc 8>>< >>>: k2\u00f0\u03b8 \u03b81\u00de \u00fe k1\u00f0\u03b81 \u03b80\u00de \u03b8 > \u03b81 k1\u00f0\u03b8 \u03b80\u00de \u03b80 < \u03b8 < \u03b81 0 \u03b80 \u2264 \u03b8 \u2264 \u03b80 k1\u00f0\u03b8 \u00fe \u03b80\u00de \u03b81 < \u03b8 < \u03b80 k2\u00f0\u03b8 \u00fe \u03b81\u00de k1\u00f0\u03b81 \u03b80\u00de \u03b8 < \u03b81 (3) where, \u03b8 is the relative angular displacement of the DMF and K\u00f0\u03b8\u00de is the nonlinear torsional elastic force function. The equation (4) can be obtained by deforming the equation (3) K\u00f0\u03b8\u00de \u00bc \u03b1k1\u03b8 \u00fe f \u00f0\u03b8\u00de (4) where, \u03b1 \u00bc \u00f0k2=k1\u00de and f \u00f0\u03b8\u00de \u00bc \u00f0\u00f0\u03b1 1\u00de=2\u00dek1\u00f0j\u03b8 \u03b81j j\u03b8 \u00fe \u03b81j\u00de \u00fe \u00f01=2\u00dek1\u00f0j\u03b8 \u03b80j j\u03b8 \u00fe \u03b80j\u00de" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000984_j.nimb.2015.02.009-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000984_j.nimb.2015.02.009-Figure2-1.png", + "caption": "Fig. 2b. Side view of central cross section of the core shows the reactor with the cold finger, 3.5 mm diameter pressure sensor, reactor valve, quartz crucible, quartz window and o-rings.", + "texts": [ + " The IR laser diode module is installed on the back plate inside the enclosure and is connected to the lens housing by an optical fibre. Fig. 1c is an exploded view of the graphitisation reactor. The core of LHF-II is machined from a solid block of stainless steel (64 64 26 mm) and internal surfaces are electro-polished. Viton o-rings are used to seal the valves, miniature pressure transducer and the 3 mm thick, 25 mm diameter quartz window, transparent to IR from the incident laser beam and from black body emission from the heated target. Fig. 2a shows a plan view of the cross section of the core, showing three valves and three ports. The three ports and the cold finger are laser welded to the block. One port is connected to the vacuum/H2 manifold and another port is connected to the CO2 sample transfer flask or breakseal cracker. The third port is spare for future development. The valves permit connection to vacuum/H2 manifold and sample CO2 and allow the reaction volume to be isolated. Fig. 2b is side view of the central cross section of the core, revealing the reaction volume and the positions of the stainless steel cold finger, crucible and quartz window. The volume of the isolated reactor is 0.25 mL, calculated by measuring the pressure of a known quantity of CO2 gas in the reactor. The catalyst is placed in a gold-plated quartz crucible (4 mm high, 4 mm diameter, 2.0 mm radius hemispherical recess) directly under the quartz window. Fig. 3 is a photograph of the vacuum manifold assembled on the work bench, also partially shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003437_s00170-020-05886-7-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003437_s00170-020-05886-7-Figure2-1.png", + "caption": "Fig. 2 Cooling channels: a conventional cooling channels and b conformal cooling channels [23, 24]", + "texts": [ + " The ability to translate a mechanism design directly to a functional mechanism manufactured from a single monolithic material represents the enormous potential for cost-savings and the predictability of subsequent performance. Flexure designs based on AM differ from existing flexure designs. First, flexures can be designed using AM process by emphasizing on the functional aspects without considering whether they can be machined. For instance, the existing cooling channel can be obtained by conventional machinability as shown in Fig. 2a. However, if it is manufactured using AM technology, forming the internal channel will not be affected by its machinability. Therefore, it can be designed as Fig. 2b without considering the limitation of machinability and by focusing only on improving its cooling performance [23, 24]. The flexure mechanisms can be designed by using AM to improve performance aspects rather than to increase possibilities of machining. When designing a conventional flexure mechanism, the machinability must be considered. Additionally, in designing a flexure mechanism, the desired properties include high precision, long feed distance, and high natural frequency. The flexure design can be performed using CAD software; however, the design geometry might not be able to be machined due to the limitation of conventional machining processes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001160_s40194-013-0107-6-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001160_s40194-013-0107-6-Figure1-1.png", + "caption": "Fig. 1 Characteristic angles of incidence at the absorbing surface in laser cutting process. The white arrow indicates the laser beam direction [9]", + "texts": [ + " Nonvertical position means that the cutting head is tilted sideways or horizontally or even in the overhead position. Nonvertical cutting has numerous applications in different areas of metal working, e.g., cutting of welding grooves for various arc welding processes. There are several manufacturers that offer systems for welding groove preparation with a CO2 laser, but only few with a fiber laser source. Some of the test results that are presented in this paper are from previous research on the subject [6\u20138]. Figure 1 shows the characteristic angles of incidence at the cut front in laser cutting. It is known that the cut front is not perpendicular to the sheet surface and therefore the actual beam incidence angle is between 75\u201390\u00b0 to the sheet top surface level [9,10]. This angle varies according to material, beam wavelength and process parameters. In conventional laser cutting of steel, the cutting head is usually pointed perpendicularly to the material surface and the cutting gas flows coaxially to the beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003437_s00170-020-05886-7-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003437_s00170-020-05886-7-Figure3-1.png", + "caption": "Fig. 3 AM-printed nanopositioning flexure mechanism [25]", + "texts": [ + " Furthermore, the ideal motion is generally designed symmetrically from one side to another to account for the thermal expansion of materials and machining errors. Therefore, the symmetrical design is structurally stable but the design results in a large and heavy system. Hardware size and weight are issues in precision equipment. In general, sophisticated technology is required to fabricate smaller systems while maintaining the desired performance. Flexure designs based on AM enable a compact system as well as maintaining the symmetry of the system. A typical example of the flexure mechanism designed using the advantages of AM is depicted in Fig. 3 [25]. While maintaining the symmetrical shape for the XY motion of the platform, the XY amplification mechanism was placed in one part of the outer section. The example flexure design as shown in Fig. 3 cannot be achieved through conventional machining; however, it can only be realized using AM. Overall, flexure mechanisms are allowed to be compact, symmetric, and complex shaped geometry by using AM technology. Because the research on the nanopositioning systems or flexure mechanisms has been of interest in academia and industry a long time ago, many patents and journals are published. Novel ideas and approaches have been historically introduced in terms of design, control, and fabrication. There exist few records for intellectual property (IP) and journal publications that are based on AM technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003678_ecce44975.2020.9235946-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003678_ecce44975.2020.9235946-Figure4-1.png", + "caption": "Fig. 4. Magnetic flux density distribution of the six models at 30,000 (rpm) for a high speed rotor, 3,000 (rpm) for a low speed rotor, the position of highspeed rotor \ud835\udf03\ud835\udf03\u210e = 90 (deg), the maximum torque output and no suspension current.", + "texts": [ + " The rotational speed of the high speed rotor is 30,000rpm, the rotational speed of the low speed rotor is 3,000rpm. The stator conductor slot fill factor is 38%. The number of coil turns is determined from the stator slot area and winding diameter of 0.4mm to keep the slot fill factor constant. The winding current was set to generate the maximum torque of the magnetic gear part. In addition, the stator core outer diameter is 130mm, the stack length is 10mm, and the sleeve thickness of the high-speed rotor is 1mm. Fig. 4 (a)-(f) shows structures and flux density distribution of the six models. Fig. 5 shows losses and efficiency of the six models. The model 1s is the initial model of this paper. The high speed rotor is equipped with a 1 mm thick sleeve. The inner air gap length is 0.3 mm, thus the total magnetic gap length is 1.3 mm. The outer gap length, that is between the low speed rotor and the stator inner PMs, is 0.3 mm. This model uses sintered magnet NdFeB N55 for both the high-speed rotor and the stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003911_s42417-020-00269-4-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003911_s42417-020-00269-4-Figure17-1.png", + "caption": "Fig. 17 Simplified scheme of clutch spring slot, with a cylindrical compression spring for preload (yellow) and a conical spring for nonlinear stiffness (blue)", + "texts": [ + " Its stiffness is equal to the linear regime of the conical spring and it is applied in the opposite direction to the conical spring mounting. The resulting function is defined by Eq.\u00a0(19). Similar to the cubic stiffness, the preloaded stiffness function starts with the torque slope equal to zero. Table\u00a03 shows the parameters chosen for the springs. Figure\u00a015 shows a comparison between the preloaded stiffness profile (red) and the original conical spring stiffness (green for linear regime and blue for non-linear regime). The conical spring design in CAD to be associated with the preloaded cylindrical spring is shown in Fig.\u00a016. Figure\u00a017 shows, conceptually, the scheme of the preload applied to a cylindrical compression spring. The conical (19) Ppl(l) = \u23a7\u23aa\u23a8\u23aa\u23a9 kl \u2212 kl = 0, 0 \u2264 l \u2264 LT \ufffd K1\u22152 \ufffd3\u22152\ufffd 1 \u2212 \ufffd 1 \u2212 2 \ufffd 1 \u2212 \ufffd 1 + K2\u2215K 2 1 \ufffd1\u22152\ufffd\ufffd1\u22152 \ufffd3 \u2212 kl, LT \u2264 l \u2264 LC. Fig. 14 Conical spring design spring, knl 12 , is activated, for instance, when the relative angular displacement is positive, which in this case is \ud835\udf032 > \ud835\udf031 . The preload force is in the opposite direction to that of the conical spring and it has increasing amplitude with the relative angular displacement ( 2 \u2212 1 ), because of the component of the linear force in the knl 12 direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure7-1.png", + "caption": "Figure 7. Motion constraints on crawler shoe 13 (Part 14).", + "texts": [ + " In Figure 6, the vectors ai, bi and ci and vectors aj, bj and cj are defined along the joint x, y and z coordinate axis defined on bodies i and j at the joint location point P. [13,15]. The two constraint equations for parallel primitive joints are from Equation (17). KP ui; uj \u00bc ci bj ai bj \u00bc 0 (17) i = 2, 3, \u2026, 13 and j = 3, 4, \u2026, 14 The propel action of the crawler track assembly is studied for two types of prescribed motion specified at the centre of mass of Part 14 (crawler shoe 13) (point P14) as shown in Figure 7. In the first motion type, the crawler track is given only translation motion, and in the second type, the crawler track is given both translation and rotational motions. The driving constraint for both types of motions is described below. In this motion, the crawler track assembly is constrained in the global positive x-direction for translation with a velocity vx(t) that varies as a cubic function of time (t) but allowed to move freely in the remaining five DOFs (two translation and three rotation)", + " D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 KTD u14; t\u00f0 \u00de \u00bc Bx14 B0;x14 Z vx(t)dt \u00bc 0 (18) B0;x14 \u00bc 17:998m is the initial global x-location of Part 14 (crawler shoe 13) shown in Figure 3 at time t = 0 and vx t\u00f0 \u00de \u00bc 0:012t2 0:0016t3 ; t1 t t2 0:1 ; t[ t2 (19) The maximum velocity of a large mining shovel is reported by Ma and Perkins [12] as 0.25 m/s. This research limits the maximum velocity of the crawler track to 0.1 m/s. In this type, the crawler shoe 13 is given prescribed translation velocity, vx(t), defined in Equation (19) along the global positive x-direction, as well as prescribed positive rotational velocity (x0 z14 ) about the z-axis of the body-fixed coordinate system shown in Figure 7. The driving constraint for the translation motion is same as Equation (18), while the rotational motion constraint is a non-holonomic constraint and can only be expressed in terms of differential of coordinates and time as Equation (20) from Shabana [16] and Schiehlen et al. [23]. KRD u14; t\u00f0 \u00de \u00bc cos h14 _w14 \u00fe _/14 x0 z14 \u00bc 0 (20) \u03c814 is the rotation angle about original z-axis; \u03b814 \u2013 rotation about the new x-axis; \u03d514\u2013 rotation about the new z-axis; x0 z14 \u2013 angular velocity about the z-axis of body- fixed coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001053_ecc.2014.6862509-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001053_ecc.2014.6862509-Figure8-1.png", + "caption": "Fig. 8. Armature equivalent circuit and power conversion of a dc drive with external excitation.", + "texts": [ + " 15$; programmer and debugging probe already on-board; just requires USB and serial (for scope) connection) Towards industrial targets (Programmable Logic Controller, PLC) there is ongoing development concerning the application to a b&r Powerpanel (PP400) using the IDE approach with the target specific Automation Studio. As a reference application the well known cart and pendulum system in gantry crane configuration is presented, see Fig. 7. As common to real world crane applications at least the equivalent length to the center of mass of the load is unknown. Instead, we assume the equivalent length of the pendulum rod l2 as unknown but constant. The cart with mass m1 is driven by a dc drive with external excitation, see Fig. 8. In the following, we assume a very small electrical time constant (\u03c4el = LA RA , with armature inductance LA and resistance RA) compared to the mechanical one. By means of system reduction, i.e, LA \u2192 0, we introduce equivalent parameters integrating the whole drive-train (drive constant km, inertia JA, transmission ratio n, gear pinion radius r, and some mechanical damping d1) into the mathematical model of the cart. Other essential parameters are explained in Tab I. m\u03031 = m1 + JA (n r )2 d\u03031 = d1 + n2k2m r2RA (1) The model equations of the nonlinear reduced system are written in the form M (q)q\u0308 = Q \u2212 C (q, q\u0307) q\u0307 \u2212 D (q, q\u0307), with the mass matrix M (q), centrifugal- and Coriolis terms C (q, q\u0307)q\u0307, dissipative terms in the vector D (q\u0307)and generalized forces Q" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000836_ecai.2013.6636196-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000836_ecai.2013.6636196-Figure1-1.png", + "caption": "Fig. 1. Magnetic circuits 3D of the alternator with claw poles", + "texts": [ + " (5) and (6) in case of a linear environment of magnetic permeability \u03bc is [5]: div (\u03bc grad \u03a6r) = div (\u03bc Hs ) (8) Relation (8) is a Poisson\u2019s equation, non-linear with the limit condition \u03a6r\u2192 0 la \u221e. Also the model did not consider eddy currents losses and their effect over the ferromagnetic properties due to heating. In this case, the calculation is reduced to a single pair of poles of the six pairs of poles in the alternator and represents one sixth of the physical model, which is limited to the borders of the infinite planes of symmetry and joins the two poles of the same polarity, Fig.1 [3],[4]. Magnetization characteristics of the stator core made possible the magnetic core laminations and forged steel claw pole for rotor, which express B\u2019s dependency of the intensity of the magnetic field H, Fig.2 and Fig.3. III. COMPUTATION OF THE 3D MAGNETIC The computation domain 3D contains a number of 87040 nodes and needs about 2 hours for solving the transient magnetic problem, Fig.4. Assuming the rotor is moving with constant speed, we can calculate the flux on a phase by integrating the radial component of the magnetic induction through the statoric coil\u2019s surface for different successive positions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure6-1.png", + "caption": "Fig. 6. The designed connecting rod.", + "texts": [ + " It has been tightened with the bump and bottom platform using M12 bolts. The connecting rod is the link that connects the bump to the crank. It has holes on both ends for mounting it on the bump and on the disc crank using M12 bolts. The dimensions proposed have been used to calculate the critical buckling load Fcr in Sec- Please cite this article as: A. Sinha, S. Mittal, A. Jakhmola et al., Green energy ge ings, https://doi.org/10.1016/j.matpr.2020.06.248 tion 2.2. It will be made of 5 mm thick IS2062 Grade E250 sheets. It is shown in Fig. 6. The disc crank is a crank in the form of a disc as shown in Fig. 7. It has a hole at a distance of 5 cm from the center for mounting the connecting rod using an M12 bolt. It is itself mounted on a shaft of outer diameter 25.4 mm. The diameter of the disc is 15 cm. A disc shaped crank has been used because it is easy to design and manufacture. It will be made of 5 mm thick IS2062 Grade E250 sheets. A compound gear train has been designed and assembled to serve as the power transmission and speed multiplication unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001099_j.jbiomech.2013.06.019-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001099_j.jbiomech.2013.06.019-Figure1-1.png", + "caption": "Fig. 1. SLIP model during human hopping. All mass is located in the center of mass (CoM) and the leg length is assumed to be the distance between CoM and center of pressure (CoP). The misalignment of ground-reaction force (GRF) and leg spring was exaggerated to illustrate the GRF contributions parallel and perpendicular to the leg direction, F\u2225 and F\u22a5 , respectively.", + "texts": [ + " Following this argument, it was shown for a one-dimensional spring\u2013mass model that simultaneous variations of leg-spring parameters (stiffness, rest length) during ground contact result in stable, robust and efficient hopping (Riese and Seyfarth, 2012a,b), motivating the variable-leg-spring (VLS) concept. So far, the results of previous studies indicate variations of leg stiffness and rest length during human hopping (Farley et al., 1991; Hobara et al., 2011), however without explicitly addressing them. Thus, in order to validate the VLS concept with experimental data, here the behavior of leg stiffness and rest length in vertical human hopping is investigated, assuming a tunable leg spring (Fig. 1). We hypothesize that the linear CoM dynamics observed in human hopping result from the interaction of non-linear leg-spring properties, namely non-constant leg stiffness and rest length, and that these parameter variations may be of considerable magnitude (410% of the touchdown value). According to Farley et al. (1991) human hopping patterns for frequencies below the preferred frequency differ from those above the preferred one. Furthermore, it was shown that humans employ different control strategies to stabilize hopping depending on the rate of movement (Yen and Chang, 2010; Hobara et al", + " In order to obtain kinematics, 17 reflective markers were placed on anatomical landmarks of each subject (Table 2). Marker positions were measured with 240 Hz using a ten-camera infrared system (Proflex MCU240, Qualisys, Gothenburg, Sweden) and interpolated to 1 kHz to match the GRF and CoP data. CoM position was then calculated in accordance with Dempster's body-segment parameter data (Dempster, 1955; Winter, 2009). 2.3. Estimation of stiffness and rest length In order to estimate global leg properties, the leg was approximated as a massless spring, connecting CoM and CoP (Fig. 1). GRFs and CoM movement during stance were projected into leg direction. Therefore, the data set is reduced to onedimensional (vertical) hopping. Additionally, GRFs were normalized to body weight (BW) and instantaneous leg length was normalized to initial CoM height linit, i.e. leg length while standing quiet. Thus, estimated stiffness K\u00bcklinit/BW and rest length L0\u00bc l0/linit are nondimensional. As the GRF and leg-length data are noisy, both data sets were filtered using a lowpass Butterworth of 5th order, with a cut-off frequency of 25 Hz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003764_j.prostr.2020.10.069-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003764_j.prostr.2020.10.069-Figure3-1.png", + "caption": "Fig. 3. A schematic of each specimen type with a cross-sectional close-up showing the filament-scale geometric features. The FNG specimen (a) has extruded filaments aligned longitudinally and no visible transverse features in the cross-sectional view. The FG specimen (b) has five equidistant manually applied transverse grooves on each face (one set of which are shown in detail on the cross-sectional close-up), which mimic those naturally present in the Z specimen (c), shown in detail in the cross-sectional close-up. Filament orientation is indicated for each specimen type.", + "texts": [ + " All specimens were stored in bags with silica desiccant to reduce any possibility of moisture absorption. Fig. 2. A schematic diagram of an F specimen (a) and a Z specimen (b) - cut from their respective boxes. Dashed lines denote the contours of the test specimen with respect to the box from which they are cut. The cross-sectional close up schematic of each specimen type demonstrates the variation in extruded filament orientation of the F (a) and Z (b) specimen, which can also be seen in the dogbone cross-section. Fig. 3. A schematic of each specimen type with a cross-sectional close-up showing the filament-scale geometric features. The FNG specimen (a) has extruded filaments aligned longitudinally and no visible transverse features in the cross-sectional view. The FG specimen (b) has five equidistant manually applied transverse grooves on each face (one set of which are shown in detail on the cross-sectional close-up), which mimic those naturally present in the Z specimen (c), shown in detail in the cross-sectional close-up" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003694_ecce44975.2020.9235404-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003694_ecce44975.2020.9235404-Figure1-1.png", + "caption": "Fig. 1. Topology of the integrated motor-compressor with continuous-skewed stator and magnets.", + "texts": [ + " The rotor of an axialflow compressor rotates to accelerate the airflow, and the diffusor diffuses the airflow to obtain a pressure increase. Axialflow compressors are normally driven by turbines or in-line motors [5]. The in-line motor-driven compressor systems have faster response time and higher efficiency over the turbinedriven compressor system. However, the in-line motor-driven system still has challenges in system compactness and motor cooling issues [6]. Recently, an integrated motor-compressor in Fig. 1 was proposed to solve these issues by integrating an FSPM machine with an axial-flow compressor. The airfoil- shaped rotor has a stagger angle with respect to the axial direction [7]-[9]. In the integrated motor-compressor, the magnets and stator are also skewed to align with the rotor and maximize the torque production capability. The integrated motor-compressor has benefits in high system compactness, self-cooling capability, and high total pressure ratio [10],[11]. The skewing techniques are widely applied in the rotor bars of the induction machines to prevent magnetic locking [12]", + " 4(a), where the stator skew angle is equal to the rotor stagger angle [8]. In this paper, the unaligned stator-rotor configurations in Fig. 4(b) are evaluated. In such configurations, the airfoilshaped rotor is skewed by the stagger angle. However, the stator and magnet are not aligned with the rotor, and the stator skew angle is not equal to the rotor stagger angle. In order to account for the effect of skew in the stator and magnets, continuous-skew and step-skew are applied to the stator and magnets, which are shown in Fig. 1, and Fig. 4(a), respectively. The manufacturing of the step-skewed structure for 5546 Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 22:42:22 UTC from IEEE Xplore. Restrictions apply. the stator and magnet is much easier to realize than the continuous-skewed structure. Fig. 5 shows the front view and top view of the stator of an integrated motor-compressor. In Fig. 5 (a), the circumferential skew angle of the stator \u03b1s is calculated in (1) where r is the stator inner radius, \u03b3 is the rotor stagger angle or stator skew angle, x is the approximate skewed pitch at stator inner radius, and le is the stack length of the stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000042_s0036023619120106-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000042_s0036023619120106-Figure4-1.png", + "caption": "Fig. 4. UV-Vis spectra at \u03c4 = 0 of (1) Nb(Cl)3TtBuPP (\u0421 = 2.45 10\u22126 mol/L) in toluene and with additions of pyridine (2) 6.2 \u00d7", + "texts": [ + " Taking into account that II was prepared by treatment of complex I with hydrochloric acid and the procedure for its isolation in solid form, we should assign to II a chemical RUSSIAN JOURNAL OF structure with a hydrogen-bonded coordinated chloride ion (Nb(Cl)3TtBuPP\u00b7\u00b7\u00b7H+\u00b7\u00b7\u00b7Cl\u2212, Scheme 1). The I and II chemosensory activity towards Py was studied in freshly distilled toluene by comparing their UV-Vis spectra before and after the addition of the base at various concentrations. The distinct optical response given by the molecules of the complexes upon Py bonding is shown in Fig. 4. It is seen that the optical response of niobium(V) porphyrins depends on their chemical structure. More pronounced changes when the Py concentration increases appear in the B-band region. There is a hypsochromic shift of a band from 419 to 429 nm and a hypsochromic shift from 551 to 523 nm in the long-wavelength range of the I spectrum. The addition of Py induces a hypsochromic shift of the \u0412-band from 426 to 420 nm and the appearance of a new weak absorption band at 554 nm in the case of compound II" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001676_amm.823.161-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001676_amm.823.161-Figure3-1.png", + "caption": "Fig. 3. a) The main stages of defining the stem. b) Final model of the prosthesis hip stem.", + "texts": [ + " The Virtual Model of the Hip Prosthesis To determine the geometric parameters of the prosthesis components made by Groupe Lepine, we used direct measurement method. Also, we identified the spatial profiles of each component. Having determined the size of geometric elements of hip prosthesis, SolidWorks [7], parameterized software for three-dimensional models, was used. Using specific commands, sketch drawing plane is obtained (Fig. 2). Using the Insert command Boss / Base Extrude Depth parameter and assigning (from left window shape-defining) the amount of 9 mm solid is obtained. On the basis form we attached another additional form. In Fig. 3, a) the main steps of defining the stem model of hip prosthesis are shown. Finally, we have obtained the stem model of hip prosthesis shown in Fig. 3, b). All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 130.237.29.138, Kungliga Tekniska Hogskolan, Stockholm, Sweden-22/02/16,19:14:10) Assembling these elements, based on simple geometric constraints, it was obtained the final model of classic hip prosthesis (Fig. 5.). The Virtual Model of the Hip Joint Prosthesis To obtain the virtual model of the prosthetic hip joint, the surgical techniques were studied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001280_s12555-012-0091-4-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001280_s12555-012-0091-4-Figure4-1.png", + "caption": "Fig. 4. The manipulator has two 6-axis F/T sensors at (a) and (b). The F/T sensor at (a) is used to detect the teaching start and end times.", + "texts": [ + " Only dummy data is removed with the DPA. Trajectory Correction based on Shape Peculiarity in Direct Teaching Manipulator 1011 We built a manipulator with the direct teaching function as shown in Fig. 3. The manipulator is guided along the outer edge of the teaching area for deburring work, which is a good example of DT application [13] because it is a tedious and repetitive job. The robot that we developed contains a 6-axis F/T sensor at the end effector to measure the contact force as shown in Fig. 4. The basic purpose of this sensor is to precisely measure the contact forces to enhance the impedance control performance during direct teaching. The F/T sensor input from Fig. 4(a) is actively used to extract the shape trajectory only. 3.1. Morphological feature points 3.1.1 Data extraction: curvature & velocity The DT trajectory should be corrected for the robot, and the original profile contains noise over a wide range. First, high-frequency noise means the robot needs a very high-speed end-effector movement capability. Second, low-frequency noise due to human teaching error causes reduced performance. The proposed STC method uses the curvature and inverse of the velocity to recognize the feature points of the DT trajectory, and the shape trajectory is corrected with the extracted features" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003720_pen.25570-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003720_pen.25570-Figure1-1.png", + "caption": "FIGURE 1 Large area projection sintering system developed in the Mechanical Department, University of South Florida (Adapted from Ref. [20]) [Color figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + "[15,19] The sintering of larger areas with a lower intensity over long exposure times could resolve some of the issues surrounding powder-based additive manufacturing techniques. Focusing thermal energy over a broad area as opposed to localized regions will lower the peak temperature associated with sintering, create a more even distribution throughout the temperature profile series, and allow more time for densification\u2014potentially increasing the mechanical properties of the final specimen. Large area projection sintering (LAPS) methods in which a projected image is used to fuse an entire layer simultaneously enables longer sintering times (Figure 1). LAPS can produce parts with greater mechanical strength than LS and high-speed sintering technologies.[17,20,21] A unique advantage of sintering the entire cross-section simultaneously is that the overall build time can be maintained despite the extended thermal exposure. The long sintering times enable the full melting of the particles without degradation of the surface, thereby contributing to an overall homogeneous microstructure not seen in LS produced parts. This results in higher density parts with greater ductility in comparison to other forms of powder-based additive manufacturing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002214_j.proeng.2016.06.231-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002214_j.proeng.2016.06.231-Figure2-1.png", + "caption": "Fig. 2. Demonstrating how diameter normal compression ratio variables were derived from HSV recorded images", + "texts": [ + " Several different deformation metrics have been reported in literature, centre of mass (COM) displacement and diameter normal compression ratio (NCR), both commonly reported metrics [9]. Diameter normal compression ratio was selected for this study, given the inability to record force-time data in this study. The diameter of the image of the ball taken immediately prior to impact ( was compared to that in the frame depicting maximal deformation during contact ( , from this the diameter normal compression ratio was calculated. (2) Deriving diameters from the HSV footage is demonstrated in Figure 2. NCR values were calculated for each ball and impact condition and averaged. Shore A hardness was selected as the hardness metric as it is used by ball manufacturers in assessing the foam layer material properties during production. In order to assess the Shore A hardness of each ball, a section of each ball was removed after all other testing, which included all material layers, since the confines of the test equipment required flat material samples. Each sample was tested five times using a Shore Scale Durometer Hardness Tester, with the mean being calculated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003813_ffe.13393-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003813_ffe.13393-Figure4-1.png", + "caption": "FIGURE 4 (A) Global model without inclusion and (B) submodel with oxide inclusion", + "texts": [ + " Therefore, the solving of global model of gears with inclusion will cost vast computation resources because there is local stress concentrate near the inclusion. In order to reduce the computing scale and ensure calculating accuracy, the submodel method is introduced in this work, which is also called the cutting boundary displacement method. The boundary condition (the displacement and stress field of gears) of submodel can be determined by the results from global model calculation. Thus, the global model of gears without inclusion and the submodel of cut gears with oxide inclusion are proposed for the fatigue life analysis, as shown in Figure 4. In this paper, the fatigue life evaluation of gears with oxide inclusion is mainly presented in domain \u03a9. The results demonstrate that the weak zone of gears is located at the zone of the inclusion. In order to reduce the calculation scale, the fatigue life evaluation is presented based on the representative \u03a9 domain. The choice of dimensions of representative \u03a9 domain can lead to the convergence results for the fatigue life of gears with oxide inclusion. For the finite element analysis of gear static contact, the intermediate gear shaft should keep stationary as a driven gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003678_ecce44975.2020.9235946-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003678_ecce44975.2020.9235946-Figure1-1.png", + "caption": "Fig. 1 Cross-sectional view of the proposed magnetic geared bearingless motor.", + "texts": [ + " In bearingless motor, two-axis active magnetic suspension is considered for a simplicity, thus, the proposed machine has a flat structure. The stator has a three-phase suspension winding and another three-phase winding for the rotating torque generation to act as a motor. This bearingless structure eliminates the mechanical bearing in high-speed rotor. The short-pitch toroidal windings are employed to shorten the coilend length, and to improve the torque density. The FEM analysis has been carried out for several prototype machine. II. PROPOSED MODEL AND THE PRINCIPLE Figure 1 shows the structure of the basic proposed machine. It consists of three parts: a bearingless high-speed rotor (HSR), low-speed rotor (LSR) and a stator core with windings. The HSR consists of permanent magnets, the number of pole-pair where \ud835\udc5d\ud835\udc5d\u210e = 1. The LSR has \ud835\udc5b\ud835\udc5b\ud835\udc59\ud835\udc59 segments of ferromagnetic pole pieces. Two sets of three-phase windings with \ud835\udc5d\ud835\udc5d\u210e and \ud835\udc5d\ud835\udc5d\u210e\u00b11 pole-pair are installed in the stator to generate the torque and the suspension force simultaneously. In Fig. 1, 2-pole windings and 4-pole windings are wound for torque and suspension force, respectively. The separated winding structure is drawn here, but combined winding having torque and suspension current is also possible. The stator is also equipped with \ud835\udc5d\ud835\udc5d\ud835\udc60\ud835\udc60 pole-pair permanent magnets. The \ud835\udc5d\ud835\udc5d\ud835\udc60\ud835\udc60 is equal to 9 in Fig. 1. The LSR modulates the permanent magnet flux. 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 278 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 17,2021 at 01:16:36 UTC from IEEE Xplore. Restrictions apply. The magnetic geared motor is a combination of a magnetic gear and a synchronous motor [7]. This combination uses modulated magnetic gears that can transmit torque in a noncontacting manner and have a high torque density. It is known that this configuration generally provides the higher torque density than that in permanent magnet (PM) motors", + " Thus, one of the best candidates is the model 6b with the highest efficiency. Fig. 6 shows the waveform of the x-axis radial force when the high-speed rotor is radially displaced by 0.3mm displacement in the x-direction. The high-speed rotor has two poles. Thus, the unbalanced pull force has a large pulsation when the magnetization direction and eccentricity are perpendicular to each other. Let us define the angle between the x-axis and the N-pole direction of the high-speed rotor as \ud835\udf03\ud835\udf03\u210e as shown in Fig. 1. The magnetic pull force is minimum at \ud835\udf03\ud835\udf03\u210e = 90, 270deg and maximum at \ud835\udf03\ud835\udf03\u210e = 0,180deg. Fig. 7 shows the relationship between the displacement and the unbalanced pull force. The magnetic pull force is a function of the rotor rotational position. Let us suppose that the rotor has touchdown clearance of 0.3mm, then active force of 6N and 13N are needed at \ud835\udf03\ud835\udf03\u210e = 90deg and 180deg, respectivity, to start the magnetic suspension. 282 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 17,2021 at 01:16:36 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure6-1.png", + "caption": "Fig. 6. Mechanical model of the Snakeboard.", + "texts": [ + " The forward propelling of a wheeled vehicle as the Waveboard can be mathematically explained through the study of its predecessor system, the Snakeboard. This section will help understand how a lateral oscillatory actuation results in the forward motion of the Snakeboard. In this way, the equations of motion of the system are derived, showing the main steps followed by Kuleshov in Ref. [5] . Resorting to the Appell\u2013Gibbs formulation of analytical mechanics, the basic principle of Snakeboard maneuvering is illustrated by imposing the time evolution of angles \u03c6r , \u03c6f and \u03c8 . For this purpose, the wellknown Snakeboard robot model shown in Fig. 6 (see Refs. [3\u20135] ) is used. In order to consider the effect of the human torso, a rotor located at the center of the mechanism is included. The mechanical model of the system consists of six coordinates: x and y locate the center of mass of the Snakeboard; \u03b8 determines its orientation in the plane; \u03c6r and \u03c6f describe the rotations of the rear and front wheelsets, respectively; and \u03c8 is needed to represent the relative rotation of the rotor with respect to the longitudinal axis. Given that the rider inputs (torques r , f and \u03c8 previously mentioned) control the rotations of the platforms \u03c6r , \u03c6f and the rotor \u03c8 , from now on it is going to be assumed that time evolutions of these coordinates are known: \u03c8 = \u03c8 ( t ) , \u03c6 f = \u03c6 f (t) , \u03c6r = \u03c6r (t) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002961_0036850419897221-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002961_0036850419897221-Figure1-1.png", + "caption": "Figure 1. Brush seal ring. (a) Brush seal ring and (b) structure of the brush.", + "texts": [ + "cn Creative Commons Non Commercial CC BY-NC: This article is distributed under the terms of the Creative Commons Attribution-NonCommercial 4.0 License (https://creativecommons.org/licenses/by-nc/4.0/) which permits non-commercial use, reproduction and distribution of the work without further permission provided the original work is attributed as specified on the SAGE and Open Access pages (https://us.sagepub.com/en-us/nam/open-access-at-sage). The specific structure of a brush seal ring consists of a front baffle, a back baffle of uniform thickness, and length bristles, as shown in Figure 1.12 Due to the nonuniformity of the interspaces between the bristles, the incoming flow becomes nonuniform and has a self-sealing effect, like comb teeth, in countless labyrinth seals during fluid flow. Because of the special seal structure of brush seals, scholars have proposed several theoretical models for numerical simulation. The uniform staggered pipe model was suggested by Kundsen in 1958,13 the square staggered column model proposed by Chupp in 1991,14 and the hexagonal staggered pipe model proposed by Holle in 1992" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003443_j.procir.2020.04.140-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003443_j.procir.2020.04.140-Figure5-1.png", + "caption": "Fig. 5. Equivalent clearance link for clearance vector model", + "texts": [ + " For this, the clearance is modeled as the virtual, massless link connecting the two centre points of the joint pair, whereby the contact surface is assumed to be rigid and friction is neglected. As friction within the joints is neglected, the direction of the clearance vector coincides with the normal direction of the collision plane. Under this assumption the clearance vector c, consisting of the clearance c and the force angle \u03b3, points in the same direction as the joint force [16, 17]. The modeling of the joint clearance and the resulting equivalent clearance link for the clearance vector model is shown in Fig. 5. The joint forces and thus the clearance vectors for moving systems can be derived using common software programs, e.g. MSC ADAMS. With the help of the MBS, the missing informations for the clearance vector like the force angle \u03b3 can be determined for a certain modeling type. Subsequently, this information can be used for solving the clearance vector loop equations and then be integrated into the tolerance analysis for a realistic representation of the joint clearance. Furthermore, lubricants in the non-assembly mechanism are not existing, so a contact load can occur, whereby the resulting impulse forces are transferred to following mechanical parts", + " (8) In order to determine the clearance vector for the enhancement of the vector loop model, as described previously, a MBS is neglected. As friction within the joints is neglected, the direction of the clearance vector coincides with the normal direction of the collision plane. Under this assumption the clearance vector c, consisting of the clearance c and the force angle \u03b3, points in the same direction as the joint force [16, 17]. The modeling of the joint clearance and the resulting equivalent clearance link for the clearance vector model is shown in Fig. 5. Fig. 5. Equivalent clearance link for clearance vector model The joint forces and thus the clearance vectors for moving systems can be derived using common software programs, e.g. MSC ADAMS. With the help of the MBS, the missing informations for the clearance vector like the force angle \u03b3 can be determined for a certain modeling type. Subsequently, this information can be used for solving the clearance vector loop equations and then be integrated into the tolerance analysis for a realistic representation of the joint clearance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002290_speedam.2016.7525820-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002290_speedam.2016.7525820-Figure11-1.png", + "caption": "Fig. 11. Stress analysis result.", + "texts": [ + " However, it is difficult to manufacture a rotor core because the structure of rotor is very complicated. Fig. 10 shows the average torque and the torque ripple of the all models. The model C has a simple rotor structure, compared with that of the model B. As a result, the flux leakage of Nd-Fe-B magnets is increased and the average torque is decreased. However, the torque performance of model C achieves the same as that of the basic model. N S N N S N S S V. ROTOR STRESS ANALYSIS We analyzed the stress in model C. The speed is 6000 min-1, which is three times the rated speed. Fig. 11 shows the Von Mises stress contour map of model C. The maximum stress is 115 MPa in the rotor core of model C. The yield strength of the steel sheet is 324 MPa and the safety factor is 3. Therefore, the Model C can be driven at 6000 min-1. CONCLUSIONS We developed the novel rotor structures for PMSM with rare earth and ferrite magnets. Model A applies the consequent pole and claw pole type motor. In this model, the total volume of the Nd-Fe-B magnets can be reduced by 57.5 % compared with that of basic model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure16-1.png", + "caption": "Fig. 16. Visualization of the SEMMT-applied part in the analysis of Motion A with Type 1.", + "texts": [ + " 14 and 15, respectively. One cycle corresponds to the period of the stroke. In Type 2, the deterioration of the mesh suddenly started around 4 cycles. The analyses of Motion A and B with Type 2 failed at 3.2 cycles and 4.5 cycles, respectively, which are described by the blue points in Figs. 14 and 15. In Type 3, the deterioration of the mesh began at early stage. The analyses with Type 3 stopped at about 3.5 cycles. On the other hand, in Type 1, the maximum aspect ratio was maintained at about 7. Fig. 16 visualizes the SEMMT-applied part in the analysis of Motion A with Type 1. It can be seen that the SEMMT-applied part deformed along with the deformation of the wing. Therefore, the shape of the SEMMT-applied part is kept as the extension of the shape of the wing. Fig. 17 shows the distribution of aspect ratio. The cross section of the SEMMT-applied part at the middle of the leading edge is shown. At different moments, the similar colored pattern can be seen, which means the mesh pattern in the SEMMT-applied part was maintained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001248_j.proeng.2014.12.005-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001248_j.proeng.2014.12.005-Figure3-1.png", + "caption": "Fig. 3. Design scheme.", + "texts": [ + " * Corresponding author. Tel.: +7-499-183-5557; fax: +7-499-183-5742. E-mail address: asv@mgsu.ru \u00a9 2014 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/3.0/). Peer-review under responsibility of organizing committee of the XXIII R-S-P seminar, Theoretical Foundation of Civil Engineering (23RSP) 2. Iterative method Consider the problem corresponding Fig. 1,a. Let the force F is applied in the core of cross-section (Fig. 3). Below is an example of calculation for the following initial data: b = 40cm, h =40cm, F = 100kN, = 5cm. We will use an iterative method. As the initial solution, we take the solution for a homogeneous material, when 7(0) 10 2.4EconstE kPa. The stresses diagram has a trapezoidal shape (Fig. 4,a). As a first approximation we consider the stress state of the column, when the modulus of elasticity varies linearly: , 2 )( 1221)1( y h EEEEyE (1) where E1 = 1.6 107kPa (the elastic modulus at y =h / 2), E2 = 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001071_s11740-013-0466-2-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001071_s11740-013-0466-2-Figure1-1.png", + "caption": "Fig. 1 Kinematic modulation in gear profile grinding", + "texts": [ + " Sintered corundum grinding wheels are used for manufacturing of case-hardened gears, which can be dressed flexibly dependent on the gear modules. The grinding process is often followed by a super finishing process, in which the surface roughness can be reduced and the material contact area on the tooth flank can be increased. Multidirectional grinding marks occur by an additional superimposed stroke movement onto the main cutting direction. An example for surface structures on the gear tooth surface is shown in Fig. 1. The chip track shape and the surface structure can be specifically changed by the E. Uhlmann C. Ba\u0308cker (&) N. Schro\u0308er Institute for Machine Tools and Factory Management, TU Berlin, Pascalstr. 8-9, 10587 Berlin, Germany e-mail: baecker@iwf.tu-berlin.de URL: http://www.iwf.tu-berlin.de kinematic modulation in abrasive machining. This process can positively influence the surface parameters and the surface structure [8]. The main objective of this research project is, to gain knowledge about the relationship between process parameters for the kinematic modulated grinding process and surface quality as well as the generation of the surface structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003842_asemd49065.2020.9276186-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003842_asemd49065.2020.9276186-Figure2-1.png", + "caption": "Figure 2. Main flux path of the AFHEM with (a) zero, (b)positive and (c) negative fields.", + "texts": [ + " MOTOR TOPOLOGIED AND PRINCIPLE OPERATION The structure of the AFHEM is shown in Fig. 1. The stator structure is composed of inner and outer stator rings, circular dc field winding, and three-phase armature winding. The dc field winding is placed between the inner and outer stators. The stator core is made of compacted SMC powder, because SMC can easily become complex shapes. The rotor is composed of two rotor disks, iron poles, permanent magnets. Figure1. Structure of the AFHEM. The principle operation of the machine is shown in Fig. 2. In the case of no field winding excitation as shown in Fig. 2(a), the PM flux travels axially from upper rotor to the lower through the stator. Fig. 2(b) shows the magnetic flux path at the positive current achieved by the dc field winding excitation, the airgap flux is reduced because the flux directions above the magnets and iron poles are opposite to each other. Fig. 2(c) shows that the direction of the flux of the iron pole is changed when the direction of the field current is changed, the flux generated by the field current at this time serves to enhance the airgap flux. III. PERFORANCE BASED ON 3D FINITE ELEMENT ANALYSIS The electromagnetic performance of the AFHEM is evaluated by 3D finite element analysis (3D-FEA). And the parameters of simulation is set as Table I. This work was supported by Natural Science Foundation of Hubei Province (Grant No. 2019CFB759) Rotor core Field Winding Outer Stator core Inner Stator core Armature Winding Permanent Magnet Iron Pole N S S SN N Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000785_s10846-013-9973-9-Figure24-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000785_s10846-013-9973-9-Figure24-1.png", + "caption": "Fig. 24 HIL Testbed schematic for the heading experiment", + "texts": [ + " 21 and 22 show that the controller successfully tracks the desired trajectory with a mean absolute error of 0.7 cm in position and 8.8 cm/s in velocity and stabilizes the lateral motion of the helicopter in the figure-8 trajectory tracking. 5.4 Heading Control Experiment To test the heading control the helicopter is mounted on the HIL testbed, shown in Fig. 23, with the arm set removed and a flat aluminum plate attached on top of the U-shape plate. The helicopter is then mounted on the flat plate such that its CG is aligned with the vertical axis of the pole as shown in the schematic in Fig. 24. Similar to the previous tests, a constant collective pitch at hover of \u03b4col = 5.6 deg is maintained after the main blades reached the nominal spinning speed of 1100 rpm. Then, the controller is activated to control the yaw motion of the helicopter to the desired trajectory. Both a step and a figure-8 yaw trajectory are tested with the results shown in Figs. 25 and 26, respectively. Figure 25 shows the yaw angle, yaw rate and tail command in the yaw control of a fast step trajectory. The results demonstrate that the controller maintains the desired yaw angle (\u03c8) and yaw rate (r) with a mean absolute error of 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002359_j.msea.2016.08.078-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002359_j.msea.2016.08.078-Figure14-1.png", + "caption": "Fig. 14. 3D microstructure model of each c", + "texts": [ + " Table 4 displays the area fractions and average equivalent grain diameters of constituent phases of the micrograph and representative Micro-CT slice and their corresponding absolute deviations. The absolute deviations of area fraction are 1.1% and 0.9% for the primary \u03b1 phase and \u03b2 phase, respectively. Correspondingly, the abstract deviations of average equivalent grain diameter are 0.2 mm and 0.07 mm, respectively. Therefore, the reliability of Micro-CT slices could be confirmed. The Micro-CT slices processed by dual-energy imaging were imported into the Simpleware to reconstruct the 3D microstructure of each constituent phase in TC6 titanium alloy. As shown in Fig. 14, the size of the 3D microstructure model is 120 100 100 mm. To analyze the 3D microstructure characteristics of TC6 titanium alloy at a finer scale, one eighth of the model was selected and the 3D models of primary and secondary \u03b1 phases as well as \u03b2 phase were displayed independently. The primary \u03b1 phase is composed of discrete equiaxial grains and interconnected grains; the secondary \u03b1 phase and \u03b2 phases form a completely interconnected network in 3D space. The reason for the 3D microstructure of the secondary \u03b1 phase appearing as an interconnected network instead of a lamellar structure is related to the heat treatment process of TC6 titanium alloy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002149_0954410016654182-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002149_0954410016654182-Figure6-1.png", + "caption": "Figure 6. Effects of the switching control gain K on the control performance.21", + "texts": [ + " The value of tanh\u00f0 s\u00de ranges from 1 to 1 with the increasing of s from negative infinity to positive infinity. Furthermore, the larger the value of , the steeper is the curve. When is close to infinity, it will degenerate to a sign function which has no benefit in reducing chattering. When is close to zero, the curve will be nearly linear and it will lose the switching property. The control law is written as u \u00bc K tanh\u00f0 s\u00de, K4 0 \u00f09\u00de where K is the switching control gain. The value of the switching control gain K has a great influence on the dynamic behavior. Figure 6 shows the effects of the switching control gain K on the control performance. The effects of different values of K on the control performance can be summarized in three directions.21 As delineated in curve \u2018\u2018a\u2019\u2019, the state trajectory reaches the sliding mode surface s \u00bc 0 in a short time for a larger value of K. However, a considerable amount of overshoot is at University of Leeds on June 19, 2016pig.sagepub.comDownloaded from noticeable. Furthermore, undesirable high control action and chattering will occur", + " The desired is shown in curve \u2018\u2018c\u2019\u2019, because the state trajectory reaches the sliding mode surface so fast, and furthermore, the chattering is almost eliminated. Figure 7 shows the block diagram of the radial basis function neural network sliding mode controller (RBFNN-SMC). It consists of two parts: a softened switching construction and a RBFNN adaptation system. The control law written as equation (9) has a fixed switching control gain. In order to achieve good tracking without causing overshoot or instability, one should minimize the reaching time and reduce chattering as shown in curve \u2018\u2018c\u2019\u2019 of Figure 6. Also, the value of the switching control gain K should be selected neither too large nor too small. Moreover, it should be adjusted online according to the states of the system. Therefore, the RBFNN adaptation system is utilized to regulate the switching control gain K online to balance the control effect and chattering phenomenon. Then, the control effect is expressed as u \u00bc K\u00fe K\u00f0 \u00de tanh\u00f0 s\u00de, K4 0 \u00f010\u00de where K is the output of the RBFNN adaptation system. Figure 8 shows the structure of the RBFNN adptation system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001301_s11665-015-1511-4-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001301_s11665-015-1511-4-Figure8-1.png", + "caption": "Fig. 8 Two-dimensional slices, at half workpiece top surface and longitudinal cross section at symmetry plane, of three-dimensional temperature field ( C) calculated using cross section information given in Table 2 for the solidification boundary (Weld 2)", + "texts": [ + " 5 and 9, it should be noted that t = 0 has been Journal of Materials Engineering and Performance assigned arbitrarily for each weld to a two-dimensional slice at the leading edge of the solidification boundary. Accordingly, shown in Fig. 5 and 9 is a passage with time of the calculated three-dimensional solidification boundaries through the experimentally measured transverse cross sections of the solidification boundary. Again, for the calculated temperature fields shown in Fig. 4 and 5, for Weld 1, and in Fig. 8 and 9 for Weld 2, the constraint condition on the calculated threedimensional solidification boundary is that projections of all its two-dimensional transverse slices, as a function of time, are consistent with the experimentally measured transverse cross section of the solidification boundary. Assuming that measurements of both the solidification and transformation boundaries are available, inverse thermal analyses of Welds 1 and 2 entail calculation of isothermal surfaces for both solidification and transformation boundaries according to experimentally measured constraint conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000395_s10409-015-0399-4-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000395_s10409-015-0399-4-Figure2-1.png", + "caption": "Fig. 2 Multi-rigid-segment model of woodpecker. A Tail. B Tarsometatarsus. C Tibia. D Femur. E Lower torso. F Upper torso. G Neck1. H Neck2. I Neck3. J Neck4. K Head. M Trunk. L Position where impact happens", + "texts": [ + " Furthermore, the tail and the tarsometatarsus (paws and connected bones) are found to be kept stationary in all cycles of pecking. To reveal the mechanism by which a woodpecker effectively attains such a high impact velocity and the roles of muscles and tendons during this process, amulti-rigid-segmentmodel of the woodpecker is established. Based on the foregoing description of the pecking process and observation of the skeletal specimen of a woodpecker, a multi-rigid-segment model, consisting of ten rods and one ball, is established (Fig. 2). In the model, the tail, tar- sometatarsus (paws and connected bones), tibia (leg), and femur (thigh) are represented by rods A, B, C, and D, respectively. The torso is represented by rods E and F. The rotation between rods E and F represents the bending of the vertebrae. The long neck is represented by rods G, H, I, and J. These ten rods are connected to their neighbors by hinges. The head is represented by a ball, K, fixed on top of the neck, J. Table 1 lists the geometry and mass of different rigid bodies in the model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003042_10402004.2020.1737285-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003042_10402004.2020.1737285-Figure5-1.png", + "caption": "Figure 5. Front view of test rig showing location of out-of-mesh thermocouple, lubricant inlet, and rotational direction.", + "texts": [ + "72- mm linear tip relief at 23 roll angle and a circular crown lead profile with nominal drop of 10.16 mm. The tests were conducted on the spur gear test rig shown in Fig. 4. It consists of two shafts arranged in parallel with two test gears (in red) and two slave gears (in green). The shaft on the left is coupled to a 75 kW electric motor (not shown) and is the driving gear. The shaft on the right is connected to a regenerative torque mechanism that employs a rotating vane system with high-pressure oil to lock torque in the loop. As shown in the front view in Fig. 5, both gears are surrounded by aluminum shrouding to direct the expelled oil and reduce windage losses. A type K thermocouple extending through the shrouding is situated just above the gear mesh to provide a close approximation of the temperature of the gear teeth by measuring the temperature of the expelled oil or air during LoL operation. During running-in, a lubricant conforming to military specification DoD-PRF-85734A for helicopter transmissions was used. This synthetic based oil contains a specially formulated additive package for enhanced load-carrying performance, antiwear, corrosion protection, and excellent thermal oxidative stability, the exact properties needed to combat the extreme operating conditions experienced inside military rotorcraft MRGBs. It is introduced to the meshing gears by a small oil jet located at the bottom of the gears as shown in Fig. 5. Preparation for testing began by subjecting each set of gears to an initial running-in period of 5 h at an operating speed of 10 krpm (pitch line velocity \u00bc 64.6 m/s). This running-in period attempts to reproduce the gear tooth surface features that exist in a transmission after operating for many hours. During the first hour, the torque in the loop was set to 23.73 N-m, producing 1.27GPa contact pressure on the teeth. The torque in the loop was then increased to 58.75 Figure 4. Spur gear test rig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000051_edpc48408.2019.9011944-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000051_edpc48408.2019.9011944-Figure11-1.png", + "caption": "Fig. 11 Symmetry conditions and wire upsetting in simulation.", + "texts": [ + " 10, the mean winding circumference can be determined with equation (2). is the inner bending radius, the core length, the core width and the wire width, respectively. , 2 \u2219 \u2219 2 8 \u2219 (2) Since increases during upsetting, while , and remain constant, resulting in an increase in , , the wire cross-sectional area cannot remain constant based on equation (1). To investigate the change in wire cross-sectional area, FEA is carried out at IBF. Due to symmetry conditions, a winding can be reduced to a quarter in the simulation. As shown in Fig. 11, a quarter winding of the coil is extracted and compressed to a final height of 0.95 mm in the simulation, starting from the initial wire cross section. The commercial software Abaqus Standard is used to perform the simulations. For the modelling of the material behaviour a Young\u2019s modulus of 130 GPa and a Poisson\u2019s ratio of 0.35 is assumed, while for the plastic material behaviour flow curves recorded at IBF are used. Since the wire length on the long side of the coil is considerably greater than the wire width, plane strain is to be expected in this area, so that the necking should occur in the corner area and the short side of the coil" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003181_tase.2020.2993277-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003181_tase.2020.2993277-Figure3-1.png", + "caption": "Fig. 3. Dynamic model of LDNSA.", + "texts": [ + " The relationship between the deflection and contact force, namely, stiffness, is no longer constant. Above all, the profile curve, where the roller moves and keeps contacting with the elastic element, determines the stiffness of the compliant mechanism. In other words, the stiffness of the compliant mechanism can be customized by designing the variation of the slope of the curve profile of the contact part. The mechanical model of the proposed LDNSA with its output shaft (inner cylinder) fixed is shown in Fig. 3 according to the structure described earlier, and the parameters are listed in Table I. Jm, bm , and Jw, bw represent the rotational inertia and damping of the motor-gearbox-encoder combination and the outer cylinder, respectively. The dynamic equations of the LDNSA system can be obtained as follows: Jm \u03b8\u0308m + bm \u03b8\u0307m = \u03c4 \u2212 \u03c4w/R (1) Jw\u03b8\u0308w + bw\u03b8\u0307w = \u03c4w \u2212 \u03c4e (2) \u03c4e = 0.15\u03b85 e \u2212 0.23\u03b84 e + 1.78\u03b83 e + 0.67\u03b8e = Ke\u03b8e (3) where \u03c4 represents the motor torque, \u03c4w is the torque of the outer cylinder that is obtained from the pulley through the wires, \u03c4e denotes the theoretical torque of the elastic components without hysteresis, and Ke is the stiffness coefficient of the elastic mechanism, which is relative to \u03b8e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001175_tmag.2014.2356593-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001175_tmag.2014.2356593-Figure8-1.png", + "caption": "Fig. 8. Electric field distribution of helicopter live-line work platform.", + "texts": [ + " In consideration of the randomness of discharge, the experiment of equipotent bonding has been implemented for six times. The arc extinguishes and reignites for several times during each experiment, and the discharge current also shows great randomness, as shown in Fig. 7, while the durations of each current pulse is between 10 and 25 \u03bcs. The current pulse 1 is chose and enlarged to demonstrate the attenuation of discharge current. A simulation model in accordance with the experimental arrangement is built and the electric field distribution of helicopter live-line work platform is shown in Fig. 8. In the simulation model, the supporting platform has been considered and treated as a conductor, whose potential is the same with helicopter live-line work platform, and the existence of the supporting platform does not affect the relation between the work platform and transmission lines. The maximum electric field on the body surface of line worker is 9.75 kV/cm. The electric field right above the head of line worker calculated by the finite element analysis is 2.16 kV/cm, which is in good agreement with the experimental result" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002634_cdc.2016.7799327-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002634_cdc.2016.7799327-Figure1-1.png", + "caption": "Fig. 1: Within the workspace Cf is the freespace flow set, a dashed line around the obstacle represents the wall following flow set Cw and obstacles are represented with crosshatching. Note the jump map is activated on the boundary of the obstacle \u2202O and resets the system a distance d away from the boundary.", + "texts": [ + " Expressing the drift and diffusion of the combined dynamics in the form of a single Fokker-Planck equation and solving it without boundary conditions, allows for an analytical solution for the systems probability density function at a desired stopping time \u03c4 . The systems\u2019 dynamics can be thought of as evolving in two distinct modes: free-space motion and wall-following. To make the discussion more concrete, examples will later be brought for W \u2282 R2. The free-space flow set is denoted Cf \u2286 W . The wallfollowing flow set Cw \u2208 F is the set of points having distance from the boundary \u2202O of the obstacle region which is in the interval [d \u2212 \u03c1, d + \u03c1], where 0 < \u03c1 < d are small positive numbers (see Fig. 1). A guard condition on the boundary \u2202O \u2208 Cf triggers a jump with map gw(x), that resets the system state to Cw. Inside the wall-following flow set Cw, a guard in a region Sl \u2208 Cw determined by a condition for disengaging from wall-following, resets the system into its free-space flow set. In this stochastic hybrid system, jumps are deterministic, flows are stochastic and for simplicity, the deterministic external forcing is removed: x \u2208Cf dx = Uf dt+Bf (x) dW (4a) x \u2208Cw dx = Uw dt+Bw(x) dW (4b) x \u2208\u2202O x+ = gw(x) (4c) x \u2208Sl x+ = gl(x) (4d) While the system is within the flow set Cf its dynamics are defined with respect to a constant reference heading \u03b8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure7-1.png", + "caption": "Fig. 7 Cutting region when peripheral cutting edge does not take part in machining", + "texts": [ + " 6 Cutting region of maximal offset emax Once offset e is bigger than emax, there will be missed cutting region, as triangular region shown in Fig. 6b. So, the maximal offset is, emax \u00bc Rc\u2212 R\u2212ap \u22c5 tan \u03b1 2 \u00f07\u00de Based on Eqs. (6) and (7), cutting depth should be calculated in two cases; 1. Peripheral cutting edge does not take part in cutting. In this case, the range of offset e is, 0\u2264e\u2264Rc\u2212 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2Rap\u2212ap2 q \u00f08\u00de 2. Peripheral cutting edge takes part in cutting. The range of offset e is, Rc\u2212 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2Rap\u2212ap2 q < e \u2264 emax \u00f09\u00de Figure 7 shows the situation that peripheral cutting edge does not take part in machining. No matter what X coordinate is, the geometrical relationship between cutting depth z and y coordinate keeps constant on plane YOZ. So we just need to figure out the expression of cutting depth z within the shaded region ABC on plane YOZ, and cutting depth of any moment can be determined by its Y coordinate. Coordinates of points A, B, C, and the center of the workpiece O1 are, yA \u00bc e\u2212 R\u2212ap tan \u03b1 2 zA \u00bc 0 ( \u00f010\u00de yB \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2\u2212 R\u2212ap 2q \u2212 R\u2212ap tan \u03b1 2 cos \u03b1 \u00fee\u2212 R\u2212ap tan \u03b1 2 zB \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2\u2212 R\u2212ap 2q \u2212 R\u2212ap tan \u03b1 2 sin \u03b1 8>>><>>>: \u00f011\u00de yc \u00bc e\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2\u2212 R\u2212ap 2q zc \u00bc 0 ( \u00f012\u00de yo1 \u00bc e zo1 \u00bc \u2212 R\u2212ap \u00f013\u00de According to the points\u2019 coordinates shown in Eqs. (10\u201313), we can obtain the mathematical expressions of boundary curve A-B-C in Fig. 7, and cutting depth of any IEi is just the z coordinates of curve A-B-C. Line A-B can be expressed as, z zA y yA \u00bc tan \u03b1 zA \u00bc 0; yA \u00bc e R ap tan \u03b1 2 ; y\u2208 yA; yB\u00bd Arc B-C can be expressed as, z zO1\u00f0 \u00de2 \u00fe y\u2010yO1\u00f0 \u00de2 \u00bc R2 zO1 \u00bc \u2212 R ap ; yO1 \u00bc e; y\u2208 yB; yC\u00bd So expressions of cutting depth z is, z \u00bc y e\u00fe R ap tan \u03b1 2 h i tan \u03b1 y\u2208 yA; yB\u00bd ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 y e\u00f0 \u00de2 q \u00fe ap\u2212R y\u2208 yB; yc\u00bd 8<: \u00f014\u00de Figure 8 shows the case of peripheral cutting edge E-C taking part in machining" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003589_1350650120962973-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003589_1350650120962973-Figure3-1.png", + "caption": "Figure 3. Schematic diagram of an average profile error (e\u00f0Mi\u00de\u00bc constant).35", + "texts": [ + " \u00bdVG\u00f0Mi\u00de \u00bc \u00bd0; 0; 0; sin\u00f0a\u00de; cos\u00f0a\u00de;Rb1; sin\u00f0a\u00de; cos\u00f0a\u00de;Rb2; 0; 0; 0 T (2) where a represents the pressure angle, Rb1 and Rb2 are the gear and pinion base radii. The contact condition for each pointMi in the base plane deduced from equation (1) is as follows: \u2022 D\u00f0Mi\u00de > 0 \u2014\u2014> Contact at point Mi \u2022 D\u00f0Mi\u00de < 0 \u2014\u2014> Lost of the contact at point Mi The error profile is specified as the distance separating the real tooth profile and the theoretical tooth profile. Two cases of interest are considered: a constant error profile is introduced in local friction coefficient formulation (see Figure 3) and a variable error profile along the action contact line and the theoretical contact line is considered (see Figure 4). The defect or shape deviation considered here is similar for all teeth. Figure 4 shows the estimated roughness profile (Rq \u00bc var) of the A10 spur gears after the end of the test in radial and axial directions with longitudinal reliefs combined to some undulations along the face width. For type A10 gear, the surface roughness in the radial orientation is generally close to the roughness in the axial orientation due to its grinding procedure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000165_ijaac.2018.092850-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000165_ijaac.2018.092850-Figure2-1.png", + "caption": "Figure 2 Two-degrees of freedom bicycle vehicle model (see online version for colours)", + "texts": [ + " On the other hand, the trailers navigation will be as following, the tail of each trailer will follow the head as a target considering the vehicle head as a goal. Also, the head of each trailers will be the tail of the previous one, Section 4.2 will illustrate the navigation function for the two trailers in details. A simplified single-track two-degree-of-freedom model (Wong, 2008) with linear tire characteristics is derived to represent the robotic vehicle motion (tractor) considering the lateral and yaw dynamics as shown in Figure 2. The forward velocity U is assumed constant and the steering angle \u03b4f is applied to the front wheel directly and assumed small (cos \u03b4f = 1). The governing equations of motion for the vehicle body lateral and yaw dynamics are given in equations (1) and (2), respectively: ( ) yf yrm V Ur F F+ = + (1) coszz yf f yrI r F \u03b4 a F b= \u2212 (2) A linear tire model is employed to estimate the tire cornering forces (Fyf, Fyr) based on the coefficients of tire cornering stiffness (C\u03b1f, C\u03b1r) and the generated tire slip angles (\u03b1f, \u03b1r) for the front and rear tires respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000404_icuas.2015.7152429-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000404_icuas.2015.7152429-Figure2-1.png", + "caption": "Fig. 2 Illustration for state and input variables", + "texts": [ + " In section III, a rigorous H\u221e static output feedback controller design procedure will be proposed with the pre-specified structure, then the performance of the controller will be validated by a series of frequency analysis and nonlinear time-domain simulations. An accurate flight dynamics model, which is based on the first principles modeling approach, has been developed for novel multi-tandem ducted fan vehicle. In addition to the kinematics, rigid-body dynamics and the tandem ducted fan dynamics, a very essential features, actuator dynamics is taken into consideration. The proposed nonlinear model is illustrated in Fig. 2. As shown in Fig. 2, the inertial frame attached to earth (NED frame) and the body-axis coordinate frame fixed at the center of gravity of the vehicle are both given. The transformation between the two frames can be derived by a standard yaw-pitch-roll Euler angles rotation. Thus the attitude and velocity in body-axis frame relative to the inertial frame is calculated by the following two equations: B c s s s c s c s c c s s B c s s s c c c s c s c s s c s c c (1) B 1 0 0 / / t s t s S c s s c c c (2) Where the short hand notation c\u0394=cos\u0394, s\u0394=sin\u0394, and t\u0394=tan\u0394", + " As derived in the previous section, both the kinematic and rigid-body dynamics equations are nonlinear differential equations, assumed to have only small perturbations and can therefore be linearized using a first order Taylor expansion of the nonlinear function in hovering operating point. Then a linearized model for the hover operating condition was established as in 0 in 0 in 0 in x A x B u Ed y C x (14) Where xin=[U, V, W, p, q, r, \u03d5, \u03b8, \u03c8]T is the state vector, u = [ ucol, ulat, ulon, uped ] T is control input vector, which are physically shown in Fig. 2. d is wind disturbance vector and the disturbance input matrix E defines the dynamics with body-frame velocities, i.e. E is a 9x3 matrix and is constituted from the first three columns of the state matrix A0, which is given as 0 0.0868 0 0 0.0023 0 0 0.0831 0 0 0 0.0096 0.0096 0.2248 0 0.0123 0 0.3805 0 0 0 0.0458 0 0 0 0.1356 0 0 0 0 0.0047 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 0 0 9.785 0 0 9.785 0 0 0.0011 0.4896 0.4896 0 0 0 0 0 0.0456 0 0 0 0.063 0 0 0 0 0 0 0 0 0 0 0 1 0 0 0 A In [11], a \u201cshaping filter\u2019, independently excited correlated Gauss-Markov processes are applied to approximate the wind components dynamics: c * c c 1 0 0 0 1 0 0 0 1 U U U V V V W WW d d q d d q d qd (15) Where \u03c4c = 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure6.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure6.6-1.png", + "caption": "Figure 6.6 A line-on-plane joint.", + "texts": [], + "surrounding_texts": [ + "A plane-on-plane joint is illustrated in Figure 6.5. It is also known as a planar contact joint. It is called so because the kinematic elements ab and ba are in contact with each other on their i \u2013 j planes. As such, its mobility is \ud835\udf07ab = 3 and it is characterized by the following orientation and location equations, which are written in terms of the joint variables xab, yab, and \ud835\udf03ab. C\u0302(ab,ba) = eu\u0303k\ud835\udf03ab (6.25) r(ab) ab,ba = uixab + ujyab (6.26)" + ] + }, + { + "image_filename": "designv11_34_0000836_ecai.2013.6636196-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000836_ecai.2013.6636196-Figure5-1.png", + "caption": "Fig. 5.Surface Sb, for calculating the flux of a coil with rotor moved with \u03b8 = 300 at t = 0,909 ms ;Ie = 2A.", + "texts": [ + " COMPUTATION OF THE 3D MAGNETIC The computation domain 3D contains a number of 87040 nodes and needs about 2 hours for solving the transient magnetic problem, Fig.4. Assuming the rotor is moving with constant speed, we can calculate the flux on a phase by integrating the radial component of the magnetic induction through the statoric coil\u2019s surface for different successive positions. This is done using \u201eTRANSIENT MAGNETIC\u201d from the program \u201eFLUX 3D\u201d. The magnetic induction mapping at \u03b8 = 300 and \u03b8 = 600, for an excitation current of Ie = 2A is represented in Fig.5 and Fig.6. The surface (Sb ) is positioned in the middle of the statoric nick having the opening equal with the coil\u2019s stepping through, in which we can calculate the flux of a coil. Induced voltage on one phase is calculated by applying the electromagnetic induction law eq. 9, taking into account N=11 turns on coil and 12 coils in series on phase. The solutions for the solved field problem and for the excitation currents Ie = 0.5; 2,5A; are represented in MATLAB in Fig.7 and Fig.8. \u2126 d\u03b8 d d d\u03b8 d\u03b8 d d d \u03a6=\u03a6=\u03a6= tt U f (9) The effective value of the voltage induced is determined with the RMS functions from MATLAB with (10) and it is represented in the table I for n=5488rot/min: Ie = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002633_icmic.2016.7804249-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002633_icmic.2016.7804249-Figure1-1.png", + "caption": "Fig. 1. Quadrotor configuration", + "texts": [ + " Section III gives an Overview of the BA and presents the PD controller design method using BA. In section IV, first the PD decentralized control structure for the stabilization of the Quadrotor is presented then BA design method is used and simulation results are presented. Section V concludes the paper. II. THE QUADROTOR MODEL The Quadrotor is an UAV minimized with four propellers mounted on the end of two perpendicular axex and actuated by four DC motors. A basic configuration of Quadrotors is shown in Fig. 1. The four rotors which form two pairs rotate inversely, one clockwise and the other counter clockwise. We consider the dynamical state model of a Quadrotor in [2] given by : 978-0-9567157-6-0 \u00a9 IEEE 2016 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0 \u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0\u00a0 (1) Where: \u2022 =\u00a0 and = are respectively the roll angle and corresponding angular velocity; \u2022 and \u00a0are respectively the roll angle and corresponding angular velocity; \u2022 and = are respectively the roll angle and corresponding angular velocity; \u2022 \u00a0, ,\u00a0and are the cartesian position coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000693_gt2014-25904-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000693_gt2014-25904-Figure5-1.png", + "caption": "Figure 5: A TYPICAL MODAL TESTS SETUP FOR A BLADE", + "texts": [ + " Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/16/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2014 by ASME Modal testing has been conducted on the test rig for the sixteen long and short blades in order to determine blade natural frequencies using the Impulse-Response modal test [7]. Each blade was excited using an instrumented hammer (Type PCB-0860C03) and the vibration responses were measured using a tiny accelerometer (Model 325C22, M/s PCB) weighing just 0.5 g, Figure 5 shows a typical modal testing setup for a blade. The blades natural frequencies were identified using the frequency response functions (FRF) calculated from the measured force and acceleration data. The experimentally identified modes for each of the long and short blades first natural frequency recorded are listed in Table 2. A small deviation in blades natural frequencies shows the presence of the blade mistuned effect due to the blade manufacturing and/or fitting. Also the natural frequencies of cracked blades are measured and found to be 121" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003430_lra.2020.3015463-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003430_lra.2020.3015463-Figure3-1.png", + "caption": "Fig. 3. Motion strategy of sub-tracks during climbing up the stair. First, main body is lifted on the stair using front side flippers. Second, sub-tracks are fully extended during climbing up the stair. Finally, front side sub-tracks are lowered to keep contact to the ground when the robot reach the upper floor.", + "texts": [ + " Trade-offs exist between the larger rotational moment and slippage for the main-tracks. Hence, a good balance between the left and right main track motions must be achieved. In this section, we describe the control rules of the main-tracks that generate rotational motions and minimize slippage. Furthermore, we describe the control rules of the sub-tracks that prevent the inclination of the main body and reduce slippages. For a full automation in climbing the spiral stairs, not only climbing, but also entering and reaching are required, as shown in Fig. 3. In the phase of \u201cEntering\u201d and \u201cReaching,\u201d a sub-track control method in [7] is used, because the requirement for the motion is the same as that in the straight stairs. In \u201cClimbing\u201d phase, We need to consider not only keeping the posture but also how to obtain rotational moment using the wall reaction force. Hence, the detailed discussion of the motion in \u201cClimbing\u201d phase is described in Section III-C. Here, the movement of the main-track is described for a clockwise climbing spiral staircase", + " The authors assumed that this was caused by the phase difference of Authorized licensed use limited to: Carleton University. Downloaded on August 23,2020 at 11:38:27 UTC from IEEE Xplore. Restrictions apply. the grousers on the left and right tracks. Based on the comparison results, the straightforward motion was discovered to be suitable for the main tracks to minimize slippages during collision. Based on the results, we set the same velocity for the left and right main-tracks to use the wall reaction force to climb up the spiral stairs. Motion of sub-tracks in \u201cClimbing\u201d phase in Fig. 3 is discussed considering how to keep the robot\u2019s posture and how to obtain rotational moment using the wall reaction force. Here, the flipper control in \u201cClimbing\u201d phase is discussed from the perspective of 1) obtaining rotational moment and 2) preventing catches with the wall or the ground. 1) Obtaining Rotational Moment: Fig. 6 shows rotational moment obtained from the wall reaction force. This figure considers three cases: (a) sub-tracks are kept flat, (b) sub-tracks are lowered, and (c) sub-tracks are raised", + " Considering the above two perspectives, sub-tracks should be kept flat during spiral stair climbing. Based on Section III-B and Section III-C, we propose a motion control method on the spiral stair. Based on the results of Section III-B, straight-forwarding motion is commanded to main-tracks. Same velocity is commanded to left and right main-tracks during climbing up the stairs. In the experiment in Section V and Section VI, commanded velocity was 0.10 [m/s]. Fig. 8 shows our proposed control algorithm for sub-tracks. This algorithm switches control rules depending on motion phases in Fig. 3 by changing the angle limitations depending on the phase. When the robot enters the stair or reaches the next floor, sub-track angles are not limited and the control algorithm works as same as [7], because the wall reaction force is not needed in these phases. In this mode, the sub-tracks are controlled to keep contact to the ground and keep the robot\u2019s posture using ground shape information measured by the LIDAR sensor and robot\u2019s inclination measured by the IMU sensor. When the robot is climbing up the stair, sub-track angle limitation works and all the sub-tracks are kept flat, considering the discussions in Section III-C", + "1 s. Fig. 15 shows the pitch inclination during an autonomous climbing trial. As shown in the figure, the maximum inclination angle was 53\u25e6. Fig. 16 shows the time series data of the angle of the front-right sub-track during autonomous spiral stair climbing. The blue line shows the results of our proposed method. The sub-track was raised when the robot entered the stairs, remained flat during climbing, and lowered when it arrived at the upper floor. This result shows that the strategy shown in Fig. 3 was realized. For comparison, the yellow line shows the result of the WITHOUT subtrack angle limitation originally introduced herein. In this case, the robot was stacked on the stair in all trials because the sub-tracks were lowered on the stairs. Fig. 17 shows a comparison of the operation time. As shown, the operation time of the proposed method was faster than those of operators B and C. The operation time of operator A was slightly more than that of the proposed method. The results show that the proposed method successfully realized autonomous spiral stair climbing, and that its operation time was faster than those of the two operators in the experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003802_s12541-020-00438-1-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003802_s12541-020-00438-1-Figure1-1.png", + "caption": "Fig. 1 Photograph and joint structure of SUBO-1 [13]", + "texts": [ + " In addition, additional constraints were added to the QP optimization formula to prevent the local ZMP from leaving the sole of the foot. In order to quickly realize the derived forces and moments of the foot, the feed forward control and the feedback control using the 6-axis force/torque sensor of the foot were simultaneously performed to improve the speed and accuracy of control. On the other hand, the Task-Space PD control law was used for controlling the horizontal position and yaw orientation of the foot. For this study, a biped walking robot, SUBO-1 was developed (Fig.\u00a01). Initially, its biped walking algorithm was studied using position control [13], however, the force-position hybrid control is proposed in this paper to improve the balance performance. Table\u00a01 shows the overall specifications of SUBO-1. The mass and height of the robot are 21.055\u00a0kg and 0.945\u00a0m, respectively. Each leg has 6 joints in total: a hip roll, hip pitch, hip yaw, knee pitch, ankle pitch, and ankle roll joints. Dimensions of the lower body were determined based on human body information" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003694_ecce44975.2020.9235404-Figure19-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003694_ecce44975.2020.9235404-Figure19-1.png", + "caption": "Fig. 19. Flux density distributions of the integrated motor-compressor with unaligned stator-rotor configurations with respect to the stator-rotor skew angles at no-load conditions.", + "texts": [ + " The magnet losses are reduced by using step-skewed segmented magnets instead of a whole piece of continuous-skewed magnet. Fig. 18 shows the magnet eddy current loss density distributions comparisons between a continuous-skewed magnet and the step-skewed magnets. The 10-segmented step-skewed magnet reduces the magnet loss by over 90%. In order to evaluate the effect of skew on the magnet and iron losses of the unaligned stator-rotor configurations, 0-degree, 8- degree, 18-degree, 28-degree, and 38-degree stator-rotor skew angles are modeled and calculated by 3-D FEA. Fig. 19 shows the flux density distributions of the five unaligned stator-rotor integrated machines with different stator-rotor skew angles but the same 38-degree rotor stagger angles at no-load conditions. Fig. 20 shows the comparisons of magnet losses, stator, and rotor iron losses with respect to the stator-rotor skew angle of unaligned stator-rotor configurations. With the increase of the stator-rotor skew angle, the flux densities in the stator cores and magnets become more evenly distributed. Therefore, the stator iron losses and magnet losses decrease with the increase of the stator-rotor skew angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001626_j.proeng.2015.12.125-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001626_j.proeng.2015.12.125-Figure2-1.png", + "caption": "Fig. 2. Spur gear. Fig. 3. Helical gear.", + "texts": [ + " Peer-review under responsibility of the organizing committee of the International Conference on Industrial Engineering (ICIE-2015) generating circuit [1]. The developed program allows for the processing of cylindrical gears with different parameters of the original contour ( , ha *, hf *, c *, *) and different coefficients of x bias generating circuit rack (Fig. 1). For the further construction of three-dimensional model using a program Autodesk Inventor, with which obtained on the basis of the profile, you can create models of various gears. To create the model spur gear (Fig. 2) is sufficient to use the extrusion operation. In contrast to the spur gear created by standard means of Autodesk Inventor (via the Design Wizard gears) profile tooth surfaces completely fit the profile, really get at run-rack instrument to the settings of the original path. In the case of a helical gear (Fig. 3) extrusion is not a straight line, and the spiral. You can use the application spring. In this application, given shape in the form of a sketch prepared tooth profile and its position relative to the axis of the wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003320_j.cmpb.2020.105646-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003320_j.cmpb.2020.105646-Figure2-1.png", + "caption": "Fig. 2. Szentpetery\u2019s digital articulator [4] .", + "texts": [ + " In 1895, most researchers considered that there is a condylar path with an inclination and developed average articulators [2] which make the condylar ball move with a fixed inclination. In 1899, Snow invented facebow to measure the information about human\u2019s skull and transfer that into the articulator to make sure the plasters of the upper jaw and the lower jaw are mounted in the correct position [3] . In 1910, Gysi developed an adaptable articulator with adjustable condylar guidance [3] . In 1997, Szentpetery [4] created a simplified but perhaps the first digital articulator to simulate digitized occlusal surfaces for articulation (see Fig. 2 ). In 2002, G\u00e4rtner and Kordass [5] recorded the jaw movement by a jaw motion analyzer (JMA) and used it to simulate the static and dynamic occlusion contacts in the computer system DentCAM. Nowadays, there are commercial digital articulators such as Ceramill Map 300 CAD/CAM system by Amann Girrbach [6] and Dental CAD/CAM System by 3Shape [7] and exocad articulator by exocad GmbH [8] . However, most commercial systems today do not have verification methods to show the accuracy of the simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001437_s0263574714001398-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001437_s0263574714001398-Figure2-1.png", + "caption": "Fig. 2. (Colour online) (a) General model of the ST robot with fixed XY and rigid body coordinate systems at time = 0; (b) The ST robot at a deformed position.", + "texts": [ + "org Downloaded: 16 Dec 2014 IP address: 138.251.14.35 three beams are flexible. Therefore, the moving platform of this robot is flexible. Additionally, it is assumed that the three beams have the same physical and geometrical properties. The angle between each two branches of the star is assumed to be 120\u25e6. Each of the three beams is joined to the rigid triangular base using a group of PRP joints. Finally, the rigid base is assumed to be an equilateral triangle. A general model of the ST robot in its start configuration is shown in Fig. 2(a). At the start of motion, the robot has no deformation. Center of the moving star and the center of the fixed triangular base coincide. Each of the three beams of the moving star intersects the corresponding side of the fixed base at its midpoint. To obtain an analytical model for vibration analysis, the three branches of the moving platform are considered. The three branches, also referred to as beams 1, 2, and 3, are each modeled as a discrete Euler\u2013Bernoulli beam with a prismatic joint. For beams 1, 2, and 3, a rigid body coordinate system is considered as x1w1, x2w2, and x3w3, respectively. See Figs. 2(a) and 2(b). The three rigid body coordinate systems are each attached to the rigid configuration of the corresponding beam. Center of the moving star at its rigid and deformed configurations are called G and G\u2032, respectively. Origins of the three coordinate systems are located at point G. The direction of each x-axis is along its corresponding rigid beam and passes through its revolute joint. Additionally, as shown in Fig. 2, a fixed coordinate system, XY with its origin O at the center of the equilateral triangle is defined. In this paper, in order to focus on off-load behavior of the robot\u2019s working point, point G, it is assumed that the end effector has zero concentrated inertia and experiences zero external load. As stated earlier, each branch of the moving platform is assumed to be an Euler\u2013Bernoulli beam. Therefore, the effects of the shear deformation and the rotational inertia moment are not considered in the motion equation of the beam element", + " (5) Position vectors for the arc centers containing the vectors \u2212\u2192 IJ and \u2212\u2192 JK can be written as \u2212\u2192 OCIJ = \u2212\u2192 OI + [ cos(\u2212\u03c0/6) \u2212sin(\u2212\u03c0/6) sin(\u2212\u03c0/6) cos(\u2212\u03c0/6) ]\u2212\u2192 IJ / \u221a 3, \u2212\u2192 OCJK = \u2212\u2192 OJ + [ cos(\u2212\u03c0/6) \u2212sin(\u2212\u03c0/6) sin(\u2212\u03c0/6) cos(\u2212\u03c0/6) ]\u2212\u2192 JK/ \u221a 3. (6) http://journals.cambridge.org Downloaded: 16 Dec 2014 IP address: 138.251.14.35 Using the position of point G and the center of the circles, we can write \u2223\u2223\u2223\u2212\u2212\u2192OG \u2212 \u2212\u2192 OCIJ \u2223\u2223\u2223 = RIJ ,\u2223\u2223\u2223\u2212\u2212\u2192OG \u2212 \u2212\u2192 OCJK \u2223\u2223\u2223 = RJK. (7) By solving Eq. (7), the position vector \u2212\u2212\u2192 OG is obtained. Then, by taking time derivative, acceleration of point G, aG is obtained as aG = d2 dt2 (\u2212\u2212\u2192 OG ) . (8) In Fig. 2(b), parameter \u03b8 represents the angular position, orientation of the moving star and is computed as follows: \u03b8 = tan\u22121 \u239b \u239d \u22121 slope (\u2212\u2192 GI ) \u239e \u23a0 , (9) in which, slope( \u2212\u2192 GI ) represents slope of vector \u2212\u2192 GI . Therefore, using Eq. (9), at the start configuration of the robot, angle \u03b8 is zero. By taking the first and second time derivatives of Eq. (9), angular velocity and acceleration of rigid motion of the moving platform can be obtained. As shown in Fig. 2(b), length of vectors \u2212\u2192 GI , \u2212\u2192 GJ , and \u2212\u2212\u2192 GK represent positions of the passive prismatic joints in the rigid body coordinate systems, x1w1, x2w2, and x3w3, as xp1, xp2, and xp3 respectively. Because the vibration amplitude and the slope angle of deformation are small, the distance between points G and G\u2032 is very small compared with the length of vectors \u2212\u2192 GI , \u2212\u2192 GJ , and \u2212\u2212\u2192 GK . Therefore, the lengths of vectors \u2212\u2192 G\u2032I , \u2212\u2212\u2192 G\u2032J , and \u2212\u2212\u2192 G\u2032K are assumed to be equal to the length of vectors \u2212\u2192 GI , \u2212\u2192 GJ , and \u2212\u2212\u2192 GK ", + " Therefore, for both the rigid as well as the flexible moving platform, the angles between the lines starting at point G and tangent to the three beams remain at the constant value of 120\u25e6. Therefore, two additional constraint equations can be written as ( \u2202w1 \u2202x1 ) G = ( \u2202w2 \u2202x2 ) G = ( \u2202w3 \u2202x3 ) G . (17) The three beams of the moving platform are assumed to be axially rigid. As stated in Section 3, due to rigid body motion of the robot, the rigid acceleration of each element of the beam in direction xi is defined as axi (Eq. (11)). Consider Fig. 2(b). Vibration of the moving platform causes point G to reach to point G\u2032. This elastic displacement causes additional axial motion of each beam which must be added to the rigid axial displacement xiG. Therefore, axial acceleration of each element of the moving platform has an additional component due to the vibrating motion of the moving platform. This introduces few nonlinear terms in the motion equations which will be further discussed in later sections. From the three equations of Eq. (14), expressions cos(\u03c8), sin(\u03c8), and wG are obtained in terms of w1G, w2G, and w3G and substituted into Eq", + " In addition, the solution provides the constraint forces/moments or the Lagrange\u2019s multipliers. In this section, four case studies, each with a different type of motion, are considered as follows: Case study 1: In place rotation s1 (t) = s2 (t) = s3 (t) = t2 Case study 2: Rectilinear translation s1 (t) = t2, s2 (t) = \u22120.5t2, s3 (t) = \u22120.5t2 Case study 3: General plane motion s1 (t) = t2, s2 (t) = 0, s3 (t) = 0.5t2 Case study 4: A specific rectilinear translation s1 (t) = t2, s2 (t) = \u2212t2, s3 (t) = 0 In all four cases, it is assumed that the robot starts in the configuration shown in Fig. 2(a). For each of the four case studies, three different groups of mode shapes are used to solve the analytical motion equations. For point G, three different boundary conditions, pinned, slide, and free, are assumed while for the free ends of the beams, points A, B, and C, we use only the free boundary conditions. Therefore, the first group uses pinned-free mode shapes, second uses slide-free mode shapes, and third uses free-free mode shapes. The results of the analytical method are also compared with FEM results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000712_detc2013-12492-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000712_detc2013-12492-Figure8-1.png", + "caption": "Figure 8. EXAMPLES FROM CLIPBOARD+", + "texts": [ + " The engineering students were enrolled in a mechanical engineering section and advised by an Assistant Professor who was teaching the capstone course for the first time; he was also supervising five other capstone projects in his section from industry sponsors and professional clients. The business students played the role of \u201csponsor\u201d and were not enrolled in or receiving any course credit for their involvement with the project. The Clipboard+ team, as they became known, conducted the capstone project just like any other project run coordinated through the Learning Factory. Literature and patent searches were conducted to understand the IP landscape, and several patents were identified (see example in Figure 8a). Concepts were generated with sketches and then refined through CAD models using customer surveys and sponsor input (see Figure 8b). Early-stage prototypes were developed, and then rapid prototyping was used to create an alpha prototype (see Figure 8c), which was then \u201cfield tested\u201d with users. The design was refined based on user feedback and testing, and a detailed manufacturing plan was developed for the final design, which included cost estimates and a bill of materials. The team also obtained a quote from a local machine shop to produce 100 units of the design. These estimates, along with the design prototype and user feedback, became instrumental in launching a fundraising campaign on Kickstarter after the capstone design project was completed (see Figure 9)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003004_j.optlastec.2020.106222-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003004_j.optlastec.2020.106222-Figure1-1.png", + "caption": "Fig. 1. (a) Schematic of the LENS processing; (b) the as-deposited for microstructural characterization.", + "texts": [ + " The nickel-based superalloy powder with a diameter of 75\u2013150 \u03bcm was obtained by gas atomization, and then the exact chemical compositions of the powder were listed in Table 1. It should be noted that the compositions are different from other commercially available nickel-based alloys such as Inconel 718 (IN718), Inconel 625 (IN625), Inconel 100 (IN100), Rene' 41 (R41), Inconel 738LC (IN738LC), and Hastelloy X (HX). The powders were dried at 100 \u00b0C for 1 h and the stainless steel substrate was polished with SiC to remove oil and cleaned with acetone. The LENS processing is illustrated in Fig. 1: the powder flow is synchronously fed through the nozzle to the focal point of the laser beam, which then melts the powder and forms a molten pool on the substrate; the laser beam and the nozzle are moved relative to the substrate or the pre-deposited layer according to the designed path, and the formed part is finally built up layer by layer. A sample (20 \u00d7 20 \u00d7 2 mm3, took 9 min) for microstructure observation, a sample (64 \u00d7 30\u00d7 4 mm3, took 162 min) for room-temperature tensile and a sample (51 \u00d7 22 \u00d7 12 mm3, took 306 min) for high-temperature tensile were fabricated, as can be seen in Fig. 1(b), Fig. 2(a) and Fig. 2(b), respectively. The process parameters were listed in Table 2. The cross-section of the as-deposited, which is parallel to the build direction (BD = z), i.e. the xz plane, was processed by electric spark cutting to serve as samples for microstructure characterization. The cross-section was polished down to diamond spray polishing agent and etched with a mixture of 10 ml CH3CH2OH + 10 ml HCl + 1 g CuCl2 for metallographic observation and Secondary Electronic(SE) characterization by using an optical microscope (LEICA DM4M) and a field-emission scanning electron microscope (Sirion200, FEI)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.24-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.24-1.png", + "caption": "Figure 10.24 MSFK-1 of the manipulator illustrated by means of L1 and L3.", + "texts": [ + "575) Case 2 with \ud835\udf03\u22171 = \ud835\udf03\u22172 = \ud835\udf03\u22173 = \ud835\udf0b, B1k = ut 1e\u2212u\u03033\ud835\udf19 \u2032 k e\u2212u\u03032\ud835\udf0be\u2212u\u03033\ud835\udefdk eu\u03032\ud835\udf0bu3 = ut 1eu\u03033(\ud835\udefdk\u2212\ud835\udf19\u2032 k )u3 = ut 1u3 = 0 (10.576) However, while Case 2 is quite likely to occur, Case 1 is not likely to occur due to the clashing possibility of the relevant links. When MSFK-1 occurs with \ud835\udf03\u22171 = \ud835\udf03\u22172 = \ud835\udf03\u22173 = \ud835\udf0b, Eqs. (10.568)\u2013(10.573) show that the rates of all the passive joint variables become indefinite and consequently the end-effector gets out of control. Therefore, it is necessary to avoid this singularity. The pose of the manipulator in this singularity is illustrated in Figure 10.24 by showing only the first and third legs (again as if \ud835\udefd3 = \u2212\ud835\udf0b). Incidentally, when Figures (10.24) and (10.21) are compared, it is seen that MSFK-1 also happens to be the PMCP (posture mode changing pose) of the manipulator. On the other hand, Eqs. (10.572), (10.573), and (10.569) imply that the manipulator may get into a pose of MSFK-2 if c\ud835\udf19\u2032 k = 0, i.e. if \ud835\udf19\u2032 k = \u00b1\ud835\udf0b\u22152 for one or more values of the index k. In any of these singularities, which looks the same as the MSIK-2 and PSIK-2 of Lk , the lower links C\u2032 kA\u2032 k and C\u2032\u2032 k A\u2032\u2032 k get aligned with themselves and the chord C\u2032 kC\u2032\u2032 k of Lk " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003367_s11432-019-2671-y-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003367_s11432-019-2671-y-Figure2-1.png", + "caption": "Figure 2 (Color online) Coordinate systems used in the model, showing the inertial, body, and fin frames.", + "texts": [ + " The electronic control systems, including the microcontroller unit (MCU) and power management modules, are also installed in this compartment. \u2022 Tail compartment. This is mainly given over to a brushless caudal motor (Maxon EC16) with a maximum rated speed of 39400 rpm, enabling the caudal joint to oscillate at very high frequencies. To provide depth control, we need to analyze the relationship between the control signal and pitch angle. To do this, we now present a brief derivation of our full-state dynamic model, based on our previous work [24]. Figure 2 illustrates the coordinate frames used, namely the inertial frame Cg = ogxgygzg, body frame Cb = obxbybzb, and rotatable surface frames Ci = oixiyizi, where i = w, t, l, r denote the waist, flukes, left flipper, and right flipper, respectively. Next, we define the translational and angular velocities with respect to (w.r.t.) the body frame as Ub = (Ubx, Uby, Ubz) T and \u2126b = (\u2126bx,\u2126by,\u2126bz) T, respectively. In addition, we define the full velocity vector as Vb = (UT b ,\u2126 T b ) T. Then, the robotic dolphin\u2019s kinematics can be expressed as gP\u0307b = gUb = gRbUb, gR\u0307b = gRb\u2126\u0302b, (1) where gRb and gPb are the rotation matrix and position vector of Cb w" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001733_ijmic.2015.072614-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001733_ijmic.2015.072614-Figure1-1.png", + "caption": "Figure 1 Vibration test platform (see online version for colours)", + "texts": [ + " Section 3 describes two steps in diagnostics: dimensionality reduction using PCA and actual classification. In this section, the new method is applied to a realistic and compared with some other diagnostic methods. Section 4 summarises our conclusions. The advantage of our proposed method is focused on a new DM method with application based on feature extraction dataset of vibration signals. The dataset was extracted when the vibration signal was analysed. Vibration features are again extracted from the time domain signal dataset by statistical method. Figure 1 shows the experimental platform that collected the vibration signal. The dataset has been extracted after vibration analysis. Our investigation object comprises 224 records containing 13 features: mean, standard deviation, variance, skewness, kurtosis, peak value, peak-peak value, square amplitude, average amplitude, waveform index, peak index, pulse index, margin index, and label (Table 1). There are 203 healthy samples and 21 fault samples with corrosion pitting on the raceways and balls in rolling bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000051_edpc48408.2019.9011944-Figure21-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000051_edpc48408.2019.9011944-Figure21-1.png", + "caption": "Fig. 21 Left: side view of semi-finished product; Right: grinding pattern and measurement of wire cross section.", + "texts": [ + " The main influencing factor is the initial width-to-height ration of the wire, which should be large. In addition, inner bending radius, friction coefficient and fillet radius of wire cross section should be large for a homogenous wire width along the winding circumference. On the other hand, a large fillet radius leads to a slipping of the windings, as it can be seen in Fig. 8 c. Based on the simulation results shown above, the geometry of the semi-finished product is adjusted. The adjusted semi-finished geometry is shown in Fig. 21. For a more homogeneous wire width along the winding circumference, the initial width-to-height ratio of the wire cross section is increased as well as the inner bending radius from 0.5 to 2.5 mm. To avoid slippage of the windings, the fillet radius is decreased to 0.5 mm. A major disadvantage of the tool shown in Fig. 6 with regard to series production is the need to tighten the wedge clamp screws to a defined torque. Therefore, in a cooperation between Breuckmann and the IBF, the multi-stage upsetting tool was further developed taking a better series production capability into account" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003430_lra.2020.3015463-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003430_lra.2020.3015463-Figure9-1.png", + "caption": "Fig. 9. Geometrical model of tracked vehicles following curved walls.", + "texts": [ + " However, a few aspects are yet to be elucidated, e.g., why a robot can follow a wall. Here, the geometrical model shows the mechanism by which the angle between the robot and the wall converges to a specific value, if the robot follows the wall because of the geometrical constraints of the wall. This model is based on the assumption that the robot continues colliding with the wall, which satisfy the conditions for stable spiral stairs climbing motion in [1]. A geometrical model is constructed as shown in Fig. 9 to analyze the rotational motion of the robot on a spiral staircase using the wall reaction force. The parameters are listed in Table I. In this model, the pose of the robot is uniquely determined by the combination of \u03c6 and \u03b8 because of the constraint where the collision point moves on the wall surface. The global frame is fixed at the center of the spiral staircase. The local frame is fixed on the robot, where the X-axis faces the front and the Y-axis faces the left. The collision point is assumed to move freely on the wall surface because passive wheels are attached on the side surface of the robot to reduce friction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002034_ijnm.2016.076159-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002034_ijnm.2016.076159-Figure1-1.png", + "caption": "Figure 1 The geometry model and mesh generation of the cutting tool (a) cutting tool modelling (b) geometry model of cutting tool (c) mesh generation of cutting tool (see online", + "texts": [ + " The workpiece material is Inconel 718, while the cutting tool is an ultrafine particle coated cemented carbide end milling tool with two flutes (MX230, NS, Japan). The physical properties of the material are given in Table 1. Due to the cutting temperature is low in micro-milling Inconel 718; the coefficients for thermal are ignored. In order to increase efficiency by reducing element sum, only the meshes in surface layer and centre cutting region of workpiece are dense, other parts are sparse. Eight nodal point, hexagon, linearly reduced integration elements are used in this paper. The geometry model and mesh generation of cutting tool are shown in Figure 1. The Johnson-Cook (J-C) plasticity model, which is suitable for modelling cases with high strain, strain rate and strain hardening, is applied to model the workpiece material, and the J-C parameters for Inconel 718 are given in Table 2 (Xu, 2012). The workpiece cutting face and tool surface are defined respectively, and the coefficient of friction is set as 0.4. Then, the tool set and reference point set are created and rigid constraints of them are established; in order to ensure rigid body displacement take place between cutting tool and specified reference point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002681_icpeices.2016.7853157-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002681_icpeices.2016.7853157-Figure2-1.png", + "caption": "Fig, 2: TWIP", + "texts": [ + " In section IV SSOSMC using SuperTwisting algorithm accompanying a comprehensive block diagram is presented along with real-time validation on LEGO Mindstorms EV3. ResuIts are analysed and conclusion is drawn in section V. 11. SYSTEM DESCRIPTION AND PROBLEM STATEMENT A two-wheeled inverted pendulwn is buiIt using LEGO Mindstorms EV3 kit and a mathematical model is obtained using its parameters. A Lagrangian approach is used here for mathematical modelling of the plant. Lagrangian approach considers generalized coordinates and it is easier as compared to Newton-Euler method of modelling. Fig. 2a i.e. the side view of TWIP shows the body pitch angle (1jI), wheel angle of left and right wheel (e\"r ), DC motor angle (f) ) and fig. 2b i.e. the top view of nI',r the TWIP shows the body yaw angle ( ifJ ). Fig. 2 is used to obtain the motion equation of TWIP[6]. The other parameters are explained in Table. 1. The following table give the parameters of LEGO Mindstorms EV3, the parameters are obtained from the real-time experimental model by performing experiments. The detailed analysis about the method to obtain the real time parameter is shown in section IIb. TABLE I: PARAMETERS OF TWIP Table Column Head Parameter Value Units Description g 9.81 rn/sec\" Gravi ty constant m 0.023 kg Mass of the wheels R 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000318_j.finel.2015.07.008-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000318_j.finel.2015.07.008-Figure4-1.png", + "caption": "Fig. 4. Rotating rigid discs model.", + "texts": [ + " 3 is chosen to describe shafts, which has 12 degrees of freedom (DOFs) and is defined as uij \u00bc xi yi zi \u03b8xi \u03b8yi \u03b8zi xj yj zj \u03b8xj \u03b8yj \u03b8zj h iT \u00f01\u00de The dynamic equation of the beam element is described as Mb \u20acuij\u00fe\u00f0Cb\u00fe\u03a9Gb\u00de _uij\u00feKbuij \u00bc Fb \u00f02\u00de where\u03a9 is the rotating speed around axes-x; Fb is the force vector of the beam element; Mb, Cb, Gb and Kb are the mass matrix, damping matrix, gyroscopic matrix and stiffness matrix respectively, where Cb is defined as Rayleigh damping and described as Cb \u00bc \u03b1UMb\u00fe\u03b2UKb \u00f03\u00de where \u03b1 and \u03b2 are the scaling factors. The impellers, wheels and gears are regarded as rigid rotating discs (with gyroscopic effect), and simplified to be lumped mass and moment of inertia acting on the corresponding node, which is shown in Fig. 4. The rigid disc has 6 degrees of freedom, which is defined as ui \u00bc xi yi zi \u03b8xi \u03b8yi \u03b8zi h iT \u00f04\u00de The dynamic equation of the rotating rigid discs is described as Md \u20acu\u00fe\u03a9Gd _u\u00bc Fd \u00f05\u00de where\u03a9 is the rotating speed around axes-x; Fd is the force vector of the rotating rigid discs; Md and Gd are the mass matrix and gyroscopic matrix respectively and described as Md \u00bc m 0 0 0 0 0 0 m 0 0 0 0 0 0 m 0 0 0 0 0 0 Jp 0 0 0 0 0 0 Jd 0 0 0 0 0 0 Jd 2 6666666664 3 7777777775 ; Gd \u00bc 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 Jp 0 0 0 0 Jp 0 2 6666666664 3 7777777775 \u00f06\u00de where m is the mass of the disc; Jp and Jd are the polar moment of inertia and the diametral moment of inertia respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000617_1539445x.2013.831358-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000617_1539445x.2013.831358-Figure2-1.png", + "caption": "Fig. 2. A schematic of the microfluidic setup used to produce the LCE droplets. A: Outer phase composed of 70 vol% water and 30 vol% glycerol with 0.36M Tween 80. B: inner oil phase containing 0.1 wt% of the LCE precursors in chloroform. C: LCE microparticles cross-linked by UV.", + "texts": [ + " After the reaction is complete, the LC polymer is purified by precipitating the polymer in ice-cold methanol. Finally, the solvent is removed by dispersing the polymer material in benzene, freezing, and sublimating under vacuum, resulting in a whitish, powdery residue. The resulting LC polymer is used as is to prepare LCE droplets. In order to generate micron sized LCE particles, we first generate emulsion droplets of a LC polymer solution in an aqueous surfactant solution using a glass capillary microfluidic setup, shown in Fig. 2, which consists of a tapered circular microcapillary and a square capillary as described in Refs. (27\u201329). The LC polymer solution is pumped through the inner circular microcapillary tip, while the outer aqueous fluid is pumped through one end of the square capillary. We note that the glass capillaries are inert to the volatile solvents which could not be used in previous PDMS and PMMA microfluidic setups, such as those described by Huang et al. (36, 37) and Zhang (38). The use of organic solvents necessitated the use of glass capillaries, as organic solvents may swell and break microfluidic devices that are made of polymers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002133_j.ifacol.2015.11.197-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002133_j.ifacol.2015.11.197-Figure6-1.png", + "caption": "Fig. 6. Model of the pendubot", + "texts": [ + " 5, shows some structural similarities to the energy-based feedback control laws proposed in A\u030astro\u0308m and Furuta (2000); A\u030astro\u0308m et al. (2005). However, in contrast to A\u030astro\u0308m et al. (2005), SAC automatically generates regions in the pendulum state space where energy is shaped by adding positive or negative damping. Moreover, although the control saturation of SAC generates large regions of \u201cbangbang-type\u201d control, we do not see the undesirable large energy overshoots as obtained from the minimum-time swing up solution of A\u030astro\u0308m and Furuta (2000) (cf. Fig. 3). The pendubot is a two-link manipulator, see Fig. 6, with only the first link actuated. The pendubot\u2019s states are its angles and velocities, x = (\u03b81, \u03b8\u03071, \u03b82, \u03b8\u03072), and its control is the torque about the attachment point of the first link, u = \u03c41. This section uses the methods previously described to compute a switching control sequence that 2 Shorter time horizons lead to more direct \u201cpushing\u201d toward the goal and longer horizons yield behaviors similar to energy tracking. 3 As a benchmark, a typical 10 s trajectory with T = 0.5 s and 100 Hz feedback requires \u2248 150 ms to compute using JE\u0304 versus \u2248 250 ms using Jx\u0304 on a laptop with an Intel i7 processor", + " Results described in this section apply the same control constraints to the LQR controller as enforced for SAC. 5.1 Stable manifold approximation and cost formulation The inverted equilibrium, x\u0304, is a hyperbolic equilibrium of the pendubot\u2019s free dynamics that is structurally equivalent to a frictionless planar double pendulum. The equilibrium has 2-D stable and unstable manifolds (cf. Fla\u00dfkamp et al. (2014)), which are computed using GAIO as described in Section 2. Figure 7 shows the box approximation of the stable manifold. Fig. 6. Model of the pendubot \u22122 0 2 \u221210 0 10 \u22122 0 2 \u03b8 1 d/dt \u03b8 1 \u03b8 2 \u221220 \u221210 0 10 20 d/dt \u03b8 2 Fig. 7. Box approximation of the stable manifold of the pendubot\u2019s inverted equilibrium. The box coloring indicates the value of the fourth coordinate, \u03b8\u03072. Over the region depicted, states on the stable manifold can be characterized as a function of angular coordinates, S : R2 \u2192 R4, S(\u03b81, \u03b82) := (\u03b81, S1(\u03b81, \u03b82), \u03b82, S2(\u03b81, \u03b82)), where Si : R2 \u2192 R, i = 1, 2, map from angular coordinates to manifold velocities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002815_1045389x19898251-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002815_1045389x19898251-Figure5-1.png", + "caption": "Figure 5. Lifting principle of L1 and L2.", + "texts": [ + " The contraction from the heated SMA spring realizes the lifting and falling motions of different driving legs. In Figure 4, the three rows show the heating modes and state changes of L1 and L2, L5 and L6, and L3 and L4, respectively, whereas the red part represents the heated part. Based on the structural design, the gravity center is ahead of L5 and L6. The lifting of L1 and L2 depends on the shift of the gravity center; L3 and L4, and L5 and L6 lift up naturally resulting from the bending deformation of the PVC body while other legs support the robot\u2019s standing. For example, in Figure 5, to raise L1 and L2, the contraction of LS1 (tail) drives TS1 to move together. Initially, L1\u2013L4 slip outward to bear the bending deformation of the PVC body accompanied with the tilt of L5 and L6; then L1 and L2 lift up when the gravity center moves to the back of L5 and L6. There are two different methods to make driving legs fall down: one is heating recovery and the other is natural cooling. The former utilizes the contraction of another SMA spring to make the PVC body recover faster, during which the PVC body is curved as an \u2018\u2018S\u2019\u2019 shape due to residual thermal stress in SMA springs", + "5 and 1 A in our experiment, respectively. The calculated lifting height is a little lower with slower initial process, due to the neglect of the contraction of the rest part of LS1 resulting from heat conduction. In Figure 10(b), the time for lifting L3 and L4 up to 1 cm requires 10 s at 1.5 A and 12 s at 1 A in our experiment, longer than the lifting of L1 and L2. The reason is that L1 and L2 slide outward more easily in the lifting of L3 and L4, leading to larger body deformation, as shown in Figure 5. The calculated lifting height is basically consistent. The calculated lifting processes of L1 and L2, and L3 and L4 driven by LS1 are shown in Figure 11, which are similar to the experimental results in the same time span (in 8 s). Dynamic simulation analysis can predict the general lifting trend of different legs, which can be utilized for the structural design to ensure designated lifting motions by changing heating modes. The lifting and falling processes of all driving legs (L1\u2013L6) with time at different driving currents are measured in Figure 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000995_s12206-013-0808-1-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000995_s12206-013-0808-1-Figure2-1.png", + "caption": "Fig. 2. The conformal cooling channel made on a plastic injection mold die using DMT process.", + "texts": [ + " It also provides metal parts coated with wear, heat, and corrosion resistant surface that suits well for metal parts in the aerospace industry. \u2022 Tool repair and remodel: DMT now can repair and recon- figure any broken or worn out mold & die to quickly be utilized as the complete tool for production. \u2022 Conformal cooling in injection mold: The full degree of geometric freedom of DMT process enables a mold die equipped with conformal cooling channel that smoothly circulated around the mold cavity, which provides efficient thermal management of the mold die for the plastic part injected as shown in Fig. 2. For fully utilizing all these advantages and applications made by DMT process, software support is now needed with advanced functionality for the process management. Today, there exist several different CAD systems that can create product geometry for different applications. Once a computer model is generated for a physical object, it needs to be communicated to another computer system for processing. In AM community, STL polygonal facet representation has become the de facto industry standard for the transfer of data, and, unfortunately, some CAD systems using surface modelers generate bad STL files with geometric flaws" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001795_robio.2015.7418892-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001795_robio.2015.7418892-Figure4-1.png", + "caption": "Figure 4. The arc plane coordinate system.", + "texts": [ + " Second, use the interpolation algorithm of 2D plane to calculate out the position and orientation of interpolation points. Last, translate the coordinate into base coordinate system. To translate the 3D question into 2D question, we need to build the arc plane coordinate system. The circle center CO is defined as the original point, with the line from origin to starting point as X-axis and the line which is normal to the arc plane and passes through original point as Z-axis. The Y-axis is determined by right-hand rule. So the arc coordinate system is built as shown in Fig 4. Projecting the unit vector of each axis of arc coordinate system to the base coordinate system, we can get the rotation matrix B B B B C C C CR X Y Z that relates arc coordinate system to base coordinate system. Suppose the position in base coordinate system of circle center CO is B COP , the homogeneous transformation matrix can be expressed as 0 0 0 1 B B B C CO C R P T (9) Suppose that is the rotational angle from starting point S to target point E and the radius of arc is R , the arc length can be expressed as R " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002751_978-3-319-00858-5_2-Figure2.14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002751_978-3-319-00858-5_2-Figure2.14-1.png", + "caption": "Fig. 2.14 Transverse pressure gradient in Poiseuille flow of nematic LCs. a The schematic representation shows the orientation of the director relative to the flow field. The director was held fixed using a strong magnetic field. b Variation of the transverse pressure gradient normalized with the primary pressure gradient plotted as a function of \u03c6. For 5CB, \u03b71 = 0.0204 Pas and \u03b73 = 0.0326 Pas were considered", + "texts": [ + "51) which gives two solutions, stable and unstable, for \u03b8 as under: \u03b8stable = + tan\u22121 \u221a \u03b13/\u03b12 and (2.52) \u03b8unstable = \u2212 tan\u22121 \u221a \u03b13/\u03b12 (2.53) for \u03b12\u03b13 > 0 only [74]. Clearly, for \u03b12\u03b13 < 0, the solutions always lead to instability (tumbling nematics). The anisotropy of the nematic viscosity leads to distinct effects in Poiseuille flow, with no isotropic analogy. One such effect is the generation of a transverse viscous stress [77, 78] when the director is held at an angle \u03c6 to the shear plane, shown in Fig. 2.14a. The director is then represented as n = (cos \u03c6, sin \u03c6, 0). Consequently, \u03c3\u0303x,z = \u03b7e f f \u2202v \u2202z where (2.54) \u03b7e f f = \u03b71 cos2 \u03c6 + \u03b73 sin2 \u03c6 (2.55) Similarly, \u03c3\u0303y,z = [(\u03b73 \u2212 \u03b71) cos \u03c6 sin \u03c6] \u2202v \u2202z (2.56) Additionally, employing the conservation of mass and momentum and the constitutive equation of the stress tensor, we can write, 0 = \u2212\u2202 p \u2202x + \u03b7e f f \u22022v \u2202z2 and (2.57) 0 = \u2212\u2202 p \u2202y + (\u03b73 \u2212 \u03b71) cos \u03c6 sin \u03c6 \u22022v \u2202z2 (2.58) By eliminating \u22022v/\u2202z2, we obtain the ratio between the transverse and longitudinal pressure gradients: \u2202 p/\u2202y \u2202 p/\u2202x = (\u03b73 \u2212 \u03b71) cos \u03c6 sin \u03c6 \u03b71 cos2 \u03c6 + \u03b73 sin2 \u03c6 (2.59) Figure 2.14b shows the variation of this pressure gradient ratio as a function of the angle \u03c6 for 5CB in nematic phase. The transverse pressure gradient has a maximum of 24 % of the primary pressure gradient at around \u03c6 = 54\u25e6. Topological defects and flow interact with each other due to the coupling between the flow and the director field. This interaction is manifested usually in two forms: (i) Influence of topological defects on the flow field and vice versa [58] and (ii) possible generation of flow due to the interaction between topological defects [79, 80]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001110_6.2014-4161-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001110_6.2014-4161-Figure3-1.png", + "caption": "Figure 3. Attitude dynamics notation. The target is supposed to y with the long axis aligned to the orbital velocity direction.", + "texts": [], + "surrounding_texts": [ + "The dynamical model used by the planner are the linearized in-plane Schweighart-Sedwick equations12 applied to the curvilinear relative states: dw\u0303 dt = 0 0 1 0 0 0 0 1( 5c2 \u2212 2 ) \u03c92 orb 0 0 2\u03c9orbc 0 0 \u22122\u03c9orbc 0 w\u0303 + 0 0 0 1 \u2206fd (t,w, \u03b4, u) , (27) where \u03c9orb and c are the constant angular orbital velocity and the Scweighart-Sedwick coe cient de ned in Section 4.A, respectively. These equations assume a circular reference orbit, secular-only perturbations of the Earth oblateness, J2. According to the assumptions introduced in Section 4.A, we note that the di erential drag is linear in u, i.e., it can be recast in the form: \u2206fd (t,w, \u03b4, u) = \u2206fd,u (t,w, \u03b4)u+ \u2206fd,0 (t,w, \u03b4) (28) The system 27 is di erentially at. It can be veri ed that a possible at output for the system is q = Q (w\u0303, u, u\u0307, ...) = \u02d9\u0303x\u2212 2\u03c9orbc y\u0303 (5c2 \u2212 2)\u03c92 orb (29) and the corresponding mapping of the states and control are w\u0303 = W (q, q\u0307, ...) = q\u0307 q\u0308\u2212(5c2\u22122)\u03c92 orbq 2\u03c9orbc q\u0308 q(3)\u2212(5c2\u22122)\u03c92 orbq\u0307 2\u03c9orbc u = U (q, q\u0307, ...) = q(4)\u2212(c2\u22122)\u03c92 orbq\u0308 2\u03c9orbc \u2212\u2206fd,0 (t,W (q, q\u0307, ...) , \u03b4) \u2206fd,u (t,W (q, q\u0307, ...) , \u03b4) (30) The constraints impose that the control variable, u, is bounded between -1 and 1. It follows q(4)(tk)\u2212 ( c2 \u2212 2 ) \u03c9orb 2q\u0308(tk) 2\u03c9orbc \u2264 \u2206fd,u (t, q, q\u0307, ..., \u03b4) + \u2206fd,0 (t, q, q\u0307, ..., \u03b4) q(4)(tk)\u2212 ( c2 \u2212 2 ) \u03c9orb 2q\u0308(tk) 2\u03c9orbc \u2265 \u2212\u2206fd,u (t, q, q\u0307, ..., \u03b4) + \u2206fd,0 (t, q, q\u0307, ..., \u03b4) k = 1, . . . , p. (31) We note that the constraints in Equation (31) are not convex in (q, q\u0307, . . .). This is due to the fact that the simple density model depends on the relative position. However, for relatively small distances (say few hundred kilometers in-track and and few hundred meters radial), it can be safely assumed that the local density is the same for the two satellites. In this way, the right hand term of the equations becomes independent on the at outputs, and the set of constraints is convex. This assumption was made in the simulations presented in Section V. Alternatively, a conservative approach consists in building a convex inner polytope approximation of the constraints as proposed in.16 4.C. On-line compensator On-line compensation is mandatory to account for non-modeled dynamics in the control plant and uncertainties, whose e ect is the deviation of the real trajectory from the scheduled one. A model predictive control (MPC) algorithm is exploited for this purpose. At each evaluation, the on-line compensator solves a problem analogous to the maneuver planner. The only di erences are the boundary conditions, the xed horizon, and the performance index. Initial conditions are provided by the current states at the beginning of the evaluation at time t. MPC is based upon the receding horizon principle, so that the nal time is not dependent on event constraints like 10 of 19 American Institute of Aeronautics and Astronautics in Equation (19), but the horizon is xed to t + th. The computed corrected control is then applied to the plant for a time tc \u2264 th. The cost function is aimed at minimizing the divergence from the reference path: Jon\u2212line = \u222b t+th t (q(t)\u2212 q\u2217(t)) T P (q(t)\u2212 q\u2217(t)) dt (32) where P is a positive de nite matrix of user-supplied weights, q = [ q, q\u0307, q\u0308, q(3) ]T , and the symbol \u2217 indicates the reference path obtained from the solution of Problem (26). V. Numerical simulations The proposed case study consists of the rendez-vous between two satellites of the QB50 constellation.17 QB50 will be a constellation of 40 Double and 10 triple CubeSats.18 The launch is planned for 2016. The constellation will be deployed on a highly-inclined near-circular LEO and the satellites will be separated by several tens or hundreds kilometers. The QB50 requirements for the `standard 2U CubeSats'19 impose that the long axis of the CubeSat must be aligned with the orbital velocity. One of these standard CubeSats is considered to be the target. QARMAN, a 3U CubeSat of the constellation developed by the Von Karman Institute for Fluid dynamics and the University of Li\u00e8ge, will be the chaser. Both the target and the chaser are assumed to be provided with 3-axis magnetotorquers and 3 reaction wheels with spin axes aligned with the edges of the CubeSat. Quaternion feedback algorithm is exploited to follow the required attitude of the two satellites. The target is assumed to be passive, i.e., its ballistic coe cient cannot be controlled. The target's reference attitude is 3-axis stabilized in its minimum-drag con guration, i.e., with its long axis aligned toward its orbital velocity direction, vt. Di erential drag is imposed by changing the ballistic coe cient of the chaser. This is achieved by pitching the chaser about the orbital normal direction, z\u0302. Table 1 lists the input parameters of the numerical simulations. 5.A. Simulation environment The numerical simulations performed in this study are carried out in a highly detailed environment. Both attitude and orbital dynamics are propagated in their complete non-linear coupled dynamics. The orbital perturbations include aerodynamic force, a detailed gravitational eld with harmonics up to order and degree 10, solar radiation pressure and third-body perturbations of sun and moon. The external torques are due to aerodynamics and gravity gradient, and the models proposed by Wertz20 for the reaction wheels and magnetic rods are exploited. In this study, the modeling of the aerodynamic perturbation assumes thermal ow, variable accommodation of the energy, and non-zero re-emission velocity. Under these hypotheses, the three extensively-used simpli cations involved in drag modeling fall into defect. Speci cally, it is not true that the drag is the only component of the aerodynamic force, that the drag coe cient is constant, and that the drag is proportional to the surface exposed to the incoming ow. For complex satellite geometries, direct simulation 11 of 19 American Institute of Aeronautics and Astronautics Monte Carlo is arguably the only way of computing these coe cients. However, this technique is extremely computationally intensive. For simple convex geometries, semi-empirical analytic methods relying on the decomposition into elementary panels provide an accurate and computationally-e ective alternative. The semi-analytic method considered in this work is based upon the research of Sentman21 and Cook22 and upon the more recent contributions summarized in.23 This model was used in our orbital propagator to compute the aerodynamic coe cients of the satellites at every time step. The two satellites are modeled with a parallelepiped shape and the principal axes are assumed to be aligned with the edges of the parallelepiped. This is very appropriate when considering satellites with body-mounted solar arrays. The contribution to the aerodynamic force and torque of possible appendices, e.g., antennas, is neglected. The atmospheric model exploited in the propagator is NRLMSISE-00. Short-term random variations are included by adding a second-order stationary stochastic process to the total mass density. The power spectral density of the process is the one proposed by Zijlstra24 rescaled for the altitude of the maneuver. The atmosphere is assumed to co-rotate with the Earth, but thermospheric winds are neglected. The realization of the trajectories of the space weather proxies is performed by exploiting the same Gaussian cupola presented in Section 4.A. Table 2 summarizes the main features of the simulation environment and it compares them to the counterpart of the control plant discussed in Section 4.B. 5.B. Results The main purpose of these simulations is to assess the bene t of using the robust reference trajectory against a non-robust one. For this reason, a Monte Carlo analysis is carried out as follows: \u2022 a single propagation is performed from t = 0 to t = tobs. The deterministic and probabilistic model of the drag are estimated; \u2022 a nominal and a robust and reference trajectory are generated. The nominal one is generated without the scenario approach by using the only deterministic contributions to the drag, i.e., estimated ballistic 12 of 19 American Institute of Aeronautics and Astronautics \u2022 starting from t = tobs, 1000 realizations of the stochastic processes related to the space weather proxies are generated by exploiting the conditional probability of the Gaussian cupola de ned in Equation (17) given the values of the proxies for t < tobs. The on-line propagation is performed for each realization; The results obtained for the three modules of the controller are analyzed in the following." + ] + }, + { + "image_filename": "designv11_34_0001084_s12206-015-0217-8-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001084_s12206-015-0217-8-Figure10-1.png", + "caption": "Fig. 10. Three-dimensional FE models with different sectioning angles of the RPV lower head: (a) fH = 2.5o; (b) fH = 45o; and fH = 90o.", + "texts": [ + " NLGEOM means non-linear geometry considering large deformation. As shown in Fig. 9, it is found that use of C3D8R or C3D8_NLGEOM can obtain more conservative results. From Table 2, it is identified that C3D8R requires the smallest capacity and the shortest calculation time. Therefore, in the study, C3D8R was used as an optimal FE. Three-dimensional FEA requires a smaller and simpler FE model for performing efficient FEA. So, the study investigated effect of sectioning the RPV lower head on the total stresses. Fig. 10 depicts three-dimensional FE models with di- fferent sectioning angles of the RPV lower head. Fig. 11 shows variations of the total stress distributions on the inner surface lines of the down-hill side vs. the sectioning angle of the RPV lower head (fn = 23.1o, tn = 13.38 mm, ro = 19.57 mm, qw = 25o, dw-up = 18.613 mm, dw-down = 31.7 mm, ww = 8.82 mm). fH means the sectioning angle of the RPV lower head. As shown in Fig. 11, it is found that the effect is insignificant. So, in the study, the sectioning angle was established with 22" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001626_j.proeng.2015.12.125-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001626_j.proeng.2015.12.125-Figure4-1.png", + "caption": "Fig. 4. Involute-bevel gear", + "texts": [ + " In the case of a helical gear (Fig. 3) extrusion is not a straight line, and the spiral. You can use the application spring. In this application, given shape in the form of a sketch prepared tooth profile and its position relative to the axis of the wheel. The angle of the tooth wheel is given through a step helix and the length of the pitch circle. Solid models of gears with more complex geometry surfaces of the teeth, you can create a combination of the two above-described software systems. For example, involute bevel wheel (Fig. 4). Involute-bevel gear (IBG) - the wheel, sliced tool rack type (Gear comb, hob, grinding wheel) with variable displacement along the axis of the wheel tool [2]. A feature of such wheels is that in each face section profile is obtained with a certain offset coefficient which is changed in each section by the amount x = s tg / m, (1) where s \u2013 move; x \u2013 increment of bias; \u2013 taper angle IBG; m \u2013 module. Coefficient of radial clearance: (2) The coefficient of the tooth head height tool: cos cos** aat hh , (3) Socket module: cos m mt " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003051_s10409-020-00937-4-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003051_s10409-020-00937-4-Figure4-1.png", + "caption": "Fig. 4 Q\u2212h curves of the control process with x0 = 0.7. Dashed curve represents the unstable shapes after the saddle-node bifurcations if the controlled material point is under bilateral displacement control", + "texts": [ + " If h is smaller than 0, the shape of elastica is symmetrical to the shape with |h| relative to x axis, whichmeans Q\u2212h curves are symmetrical relative to the original point. As the controlled material point moves downward, four coexisting kinds of equilibrium shapes are observed which correspond to four branches. Given h, four kinds of shapes are determined by the cross points of two stiffness\u2212curvature curves which are introduced in Sect. 2.4, and the external force Q is obtained as introduced in Sect. 2.5. Varying h, load\u2212deflection curves are obtained as shown in Fig. 4. Branch 1 represents the shape that both elasticas A1C and A2C are not inverted. Branch 2 represents the shape that A2C is inverted. Branch 3 represents the shape that A1C is inverted. Branch 4 represents the shape that both A1C and A2C are inverted. Based on the symmetry feature of whole elastica, critical points are investigated only at branches 1 and 2. If the whole elastica is under force control, critical points exist at the local maximum points on both two branches, i.e., branch 1 and branch 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003656_s10015-020-00655-x-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003656_s10015-020-00655-x-Figure2-1.png", + "caption": "Fig. 2 Definition of the coordinate systems of the UAV robot", + "texts": [ + " After constructing a simplified dynamical model of the UAV robot, a computed torque controller is designed for the UAV robot and a coordinate transformation is used to simply the allocation problem. Some simulation results of the proposed UAV robot are given to demonstrate the tracking performance of arbitrary position and attitude. A simulation with wind disturbances is also added to verify the position control performance of the proposed control strategy in the final part. The overview of a prototype UAV robot developed in Xu et\u00a0al. [9] is shown in Fig.\u00a01. The definition of the coordinate systems of the prototype UAV robot is shown in Fig.\u00a02. The proposed UAV robot can be considered as a multibody system including three rigid bodies, i.e., two tiltable coaxial rotors Pi and a main body B . In each tiltable coaxial rotor, two RC servomotors are set for constructing a 360\u25e6 tilt mechanism, and one RC servomotor is also set for making another 180\u25e6 tilt mechanism, while two brushless motors are prepared for actuating a pair of DJI 1038 propellers. In the main body, it includes an aluminum pipe structured for the body frame, four passive wheels that are used on a wall or floor, four guide wheels that are available in a transition mode from a floor to a wall (or in its reverse transition mode), a battery and all other electronic devices", + "2 m, while the height h between them is 0.05 m. The principal moment of inertia of the UAV robot is defined as IBxx = IByy = 0.01kgm2 and IBzz = 0.006 kg\u00a0m2 , where the data of IBxx , IByy and IBzz are come from the 3D model built in Inventor. The propeller thrust coefficient kf = 1.95604 \u00d7 10\u22127N\u2215(rpm)2 is an average value obtained from the actual thrust experiments, in which the rotation speed of two brushless motors in a tiltable coaxial rotor is set to be the same to cancel the anti-torque. In Fig.\u00a0 2, it shows the defined coordinate systems and the variable definitions for the proposed UAV robot. The coordinate system FW \u2236 { OW;XW,YW, ZW } is the world inertial coordinate system, whereas the coordinate system FB \u2236 { OB;XB, YB, ZB } is the UAV robot body coordinate system rigidly attached to its center of mass. The rotation angles around each axis in FB are ( ) . The definition of the rotation matrix WRB from the body coordinate system FB to the world coordinate system FW is described as below: in which RZ , RY and RX are the rotations around the axis ZB , YB and XB , respectively. The coordinate systems FP1 \u2236 { OP1 ;XP1 , YP1 , ZP1 } , FP2 \u2236 { OP2 ;XP2 , YP2 , ZP2 } , tilt angles 1 , 1 , 2 , 2 and some other variables of the coaxial rotors are also shown in Fig.\u00a02. For the expression of the model in a convenient way, the coordinate system of the i-th coaxial rotor ( i = 1, 2 ) is shown in Fig.\u00a03. The tiltable coaxial rotor-fixed coordinate systems lie in a plane and are separated by the angle of (see Fig.\u00a02), which are given by FPi \u2236 {OPi ;XPi , YPi , ZPi } (see Fig.\u00a03). Tilt angles (1)WRB = RZ( )RY ( )RX( ) i and i denote the ith coaxial rotor tilt angles about XPi and YPi , respectively. The initial tilt angles i and i are set to 0. The possible change about the tilt angles of the tiltable coaxial rotor is shown in Fig.\u00a04. In each tiltable coaxial rotor, the 360\u25e6 tilt mechanism is attached to the body coordinate system so as to change the pitch angle i of the coaxial rotor in 360\u25e6 (see Fig.\u00a04a), additionally, the 180\u25e6 tilt mechanism is connected to the coaxial rotors so that the roll angle i of the rotor can be changed in 180\u25e6 (see Fig", + " The rotation matrix BRPi from FPi to FB is written as follows: where, RZ , RY and RX are the rotations around the axis ZPi , YPi and XPi , respectively. The position vector OPi of the ith tiltable coaxial rotor in FB is defined as: The angular velocity Pi and its acceleration \u0307Pi of the i-th tiltable coaxial rotor can be obtained as follows: (2)BRPi = RZ((i \u2212 1) )RY ( i ) RX ( i ) (3)BOPi = RZ((i \u2212 1) )[l 0 h]T (4) { P1 = [?\u0307?1 \ud835\udefd1 ?\u0304?2 \u2212 ?\u0304?1] T P2 = [?\u0307?2 \ud835\udefd2 ?\u0304?3 \u2212 ?\u0304?4] T (5) { \u0307P1 = [\ud835\udefc1 \ud835\udefd1 \u0307\u0304\ud835\udf142 \u2212 \u0307\u0304\ud835\udf141] T \u0307P2 = [\ud835\udefc2 \ud835\udefd2 \u0307\u0304\ud835\udf143 \u2212 \u0307\u0304\ud835\udf144] T 1 3 in which the rotation speed of each brushless motor is denoted as ?\u0304?j , j = 1 , 2, 3, 4 (see Fig.\u00a02). The anti-torque drag,i of the ith rotor is given by: where km is the propeller drag coefficient and km > 0 , whereas the thrust Tthrust,i (see Fig.\u00a02) of the ith coaxial rotor is in which kf is the propeller thrust coefficient mentioned in Sect.\u00a02.1. Referring to [10\u201313], a dynamical model for the proposed UAV robot can be derived. Using the Newton\u2013Eulerian\u2019s law, under the coordinate system of the coaxial rotors, the torque Pi produced by the tiltable coaxial rotor is reduced to where IP is the symmetric and positive definite inertia matrix of the coaxial rotor. In addition, using the Euler equations of motion, it follows that the angular velocity of the UAV robot expressed in the body coordinate system, B , is subject to where IB is the symmetric and positive definite inertia matrix of the body, B is defined as input torque, and ext refers to the unmodeled disturbance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000785_s10846-013-9973-9-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000785_s10846-013-9973-9-Figure2-1.png", + "caption": "Fig. 2 HIL Testbed schematic of front and side views", + "texts": [ + " For small unmanned helicopters, in-flight gain tuning is potentially risky and expensive due to the high manoeuvrability and internal instability properties of these vehicles which may result in a crash during the tuning process. In order to facilitate safe tuning of control gains and to test the controller against external disturbances in a controlled environment, a Hardware-In-The-Loop (HIL) testbed design for small unmanned helicopters has been developed and is shown in Fig. 1. Schematics of the testbed from the front and side views are shown in Fig. 2. This testbed is a 2-DOF system composed of a long pole to raise the helicopter off the floor to eliminate the ground effect and a headpiece. The pole is anchored to the ground. The headpiece is connected to the pole through a set of thrust roller bearings which allows for the rotation of the head with respect to the pole around the vertical Z-axis. The headpiece itself is composed of: a U-shape aluminum plate, two arms, a damping cylinder, a counter balance weight, and one stopper on each side", + " The helicopter is mounted on Arm 1 which is connected to Arm 2 through two ball bearings on the Ushape plate. This allows for the rotation of the helicopter around the horizontal X-axis. To allow independent testing of the longitudinal/lateral motion a mechanical device restricts the rotation of the headpiece around the vertical axis. Two adjustable hard stoppers restrict the rotation around the X-axis to \u00b130 deg. In the event of controller malfunction, a nonlinear passive damping system, shown from the front view in Fig. 2, prevents the helicopter from hitting the hard stops at a large angular speed. The damping system is composed of a double acting cylinder fully filled with a light oil NUTO A10. An oil line with an adjustable needle valve connects the cylinder\u2019s inlet to its outlet and makes a closed path of oil. Since the cylinder is double acting and its rod extends from both ends of the barrel, it allows for the continuous flow of oil between the chambers and does not require any oil reservoir. When testing the helicopter around hover on the testbed, the cylinder is vertical and creates a negligible damping due to the vertical mechanism geometry. As the helicopter deviates from hover the cylinder angle relative to the vertical pole increases and the damping increases in a nonlinear manner. Once the control gains are properly tuned for hover the damping cylinder is removed and the controller is tested against external disturbances. The counter weight is chosen to precisely balance the arms to mimic hover condition and the mechanism shown in Fig. 2 is designed such that the CG of the helicopter is aligned with the horizontal X-axis which is the axis of rotation in roll and pitch tests. The goal of the designed HIL simulation testbed is to safely test and tune the longitudinal, lateral and heading controller of the helicopter in real-time on the ground. In addition, it allows for testing the robustness of the controller against external disturbances. Using this testbed the helicopter can be disturbed during operation and the performance of the controller can be tuned", + " The calculated states of the helicopter are then used as the feedback signal to the controller for the next sampling time computations. This HIL is implemented in realtime at a sample rate of 50 Hz using xPC Target [28]. This HIL testbed allows for testing the helicopter not only in hover but also for smooth trajectories, such as the cruise flight, figure-8, etc. For instance, lateral control of the helicopter in a figure-8 trajectory can be implemented on the testbed by allowing the physical rotation of the vehicle only around the roll axis (the x-axis shown in Fig. 2) and then calculating the actual motion of the rest of the DOFs in the plant simulation using its 6-DOF nonlinear mathematical model. Since the physical motion of these DOFs are mechanically restricted on the testbed, and the controller is controlling all 6-DOFs of the helicopter, the generated control signal for them are not physically realized on the actual helicopter. Rather, some of these signals are applied to the simulated helicopter model in the plant simulation while the control signal corresponding to the moving DOF is physically applied to the actual helicopter", + " Although the helicopter is physically restricted to rotate only around the pitch axis in this experiment, the controller is con- trolling all 6-DOFs of the helicopter. The actual motion of the helicopter is calculated using the 6- DOF nonlinear model of the helicopter at every sampling time as shown in Fig. 3. The CG of the helicopter shown in Fig. 8 is aligned with the axis of rotation. Depending on the size of the blades and the maximum rolling and pitching moments generated the needle valve on the damping cylinder shown in Fig. 2 is ad- justed to prevent sudden movements that could damage the vehicle during tuning. Then once the desired gains are set, the damping effects are reduced by adjusting the needle valve and then by removing the cylinder to test the controller subject to the imposed external disturbances without damping. The Control Point (CP) is chosen to be 3 m above the CG on the main hub axis of the Evolution EX helicopter. Although the Control Point approach, described in Section 4, is used in this paper to tune the gains of the Sliding Mode controller, the testbed can be used for tuning any controller designed using any approach" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003586_s12206-020-0918-5-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003586_s12206-020-0918-5-Figure4-1.png", + "caption": "Fig. 4. Distribution of the maximum velocity evaluation index.", + "texts": [ + " 3(c)) and need to be compensated in real time, which may cause a certain influence on the velocity con- trol accuracy of the PM. The maximum and minimum velocity evaluation index of the PMs is used to define the output velocity performance [18], namely ( )T max Vmax= [ ] [ ]\u03bbV J J (26) ( )T min Vmin= [ ] [ ]\u03bbV J J (27) where ( )T Vmax [ ] [ ]\u03bb J J and ( )T Vmin [ ] [ ]\u03bb J J are the maximum and minimum eigenvalues of the matrix T[ ] [ ]J J , respectively. The larger the maximum and minimum velocity evaluation indices, in theory, the better the velocity performance of the PM. Fig. 4 presents the maximum velocity evaluation index of the PM with a completely constant, partially constant and inconstant Jacobian. When the actuated velocities are given, the moving platform moves in different poses. The PM with a completely constant Jacobian can always maintain a constant velocity evaluation index about 1.1547. The output maximum velocity evaluation index of the PPRU 2PPRU+ PM with a partially constant Jacobian remains approximately constant. And the output maximum velocity evaluation index of the 3 - PPRU PM with an inconstant constant Jacobian grows at an exponential rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000198_j.phpro.2011.03.023-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000198_j.phpro.2011.03.023-Figure1-1.png", + "caption": "Figure 1. Principle of LB-GMA hybrid welding with bottom closed keyhole", + "texts": [ + " So usable welds with plate thicknesses up to 18 mm in single side single pass technique can be performed with square butt and Y-shaped bevel preparations with high root face heights. With increasing plate thicknesses the adoption of double sided single * Corresponding author. Tel.: 0049 241 80 96253. E-mail address: olschok@isf.rwth-aachen.de. pass technique with laser beam GMA hybrid welding and the use of suitable weld preparations is a possible way of decreasing the welding passes normally required by arc processes, Figure 1. Process related double sided single pass weld seams with the hybrid process show a tendency to pore formation in the lower weld regions. This behavior has been observed with both laser beam and laser beam GMA hybrid welding processes with a penetration depth of more than 12 mm [2]. The main reasons for the pore formation with laser beam processes without a keyhole that is open to the bottom of the weld are the unique weld geometry with high aspect ratio between weld depth and weld width in conjunction with the process related short solidification times" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000678_robio.2014.7090419-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000678_robio.2014.7090419-Figure1-1.png", + "caption": "Fig. 1: Illustration of vehicle base with two manipulator arms with position and orientation of the vehicle body and the manipulator end effectors.", + "texts": [ + " In this paper a two-manipulator underwater vehicle in the plane is considered. The proposed approach can be followed for manipulators with any number of links. However, for simplicity of presentation we write the equations for the 3 and 2 links case. The vector \u03b7 b = [xb,yb,\u03c8b] T is the vehicle position and orientation relative to the inertial frame. Similarly, \u03b7 1 = [xee1,yee1,\u03c8ee1] T and \u03b7 2 = [xee2,yee2,\u03c8ee2] T describe the position and orientation of the manipulator end effectors in inertial frame, see Figure 1. The vehicle\u2019s velocity is defined in body-frame as \u03bd = [u,v,r]T , where u and v are the linear velocities and r is the angular rate. Equation (1) describes the relationship between the velocities in the inertial and body frame: \u03b7\u0307 b = Ri b(\u03b7 b)\u03bd = \u23a1 \u23a3cos(\u03c8b) \u2212sin(\u03c8b) 0 sin(\u03c8b) cos(\u03c8b) 0 0 0 1 \u23a4 \u23a6\u03bd (1) Furthermore, the manipulator arms have joint angles q1 = [q11,q12,q13] T and q2 = [q21,q22] T and joint angular velocities q\u03071 and q\u03072. The forward kinematics of the manipulator arms are straight forward" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure3-1.png", + "caption": "Figure 3. Overall view and a cross-sectional diagram of the gear with six conical passive rollers", + "texts": [ + " We examined three types of gears with three different shapes of passive rollers. In the following chapters, we introduce the designs and features of these gears with passive rollers. Another function of the gear with passive rollers as a worm wheel to achieve worm gearing with high energy efficiency is also introduced. 978-1-4673-5643-5/13/$31.00 \u00a92013 IEEE 1520 To achieve a spur gear with smooth sliding motion that drives the omnidirectional gear, we designed gears with passive rollers in various shapes. Schematic diagrams of the gears with passive rollers are shown in Fig. 3, 4 and 5. When designing these gears, we mainly considered the following two points: Miniaturization and simplification of the gear mechanism to reduce the size and weight of the whole omnidirectional gear system The cost of manufacturing the gear mechanism Initially, we examined the gear with conical passive rollers shown in Fig. 3. In this gear structure, the rotational axis of each passive roller is radially aligned. The rotational axes in a radial direction have the same offset in a circular direction around the center shaft of the whole gear structure. It is easy to miniaturize and reduce the weight of this mechanism. By adapting a structure including multiple layers of passive rollers with a projection outline of an involute curve, the number of passive rollers can be reduced to at least six, as shown in Fig. 3. The second gear mechanism we examined has flat passive rollers as shown in Fig. 4. In this gear structure, the rotational axis of each passive roller is aligned continuously on the circular trajectory around the center shaft of the whole gear mechanism. Each rotational axis of a passive roller is independent. This gear mechanism eliminates the need for multiple layers of passive rollers like the gear with conical passive rollers shown in the previous chapter for sliding with an omnidirectional gear with an intermittent teeth structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.25-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.25-1.png", + "caption": "Figure 9.25 Right elbowed and left elbowed poses of a Scara manipulator.", + "texts": [ + "308) If s\ud835\udf035 \u2260 0, the other equation pairs give \ud835\udf036 and \ud835\udf034 as follows without additional superfluous sign variables. \ud835\udf036 = atan2(\ud835\udf0e5c\u221723, \ud835\udf0e5c\u221713) (9.309) \ud835\udf034 = atan2(\ud835\udf0e5c\u221732,\u2212\ud835\udf0e5c\u221731) (9.310) (a) First Kind of Multiplicity The first kind of multiplicity is associated with the sign variable \ud835\udf0e2 that arises in the process of finding \ud835\udf032 by using Eqs. (9.296)\u2013(9.298). The manipulator attains the same location of the wrist point by assuming one of the two alternative poses that correspond to \ud835\udf0e2 = + 1 and \ud835\udf0e2 = \u2212 1. These poses are illustrated in Figure 9.25. The poses corresponding to \ud835\udf0e2 = + 1 and \ud835\udf0e2 = \u2212 1 are designated, respectively, as right elbowed pose and left elbowed pose. These designations are justified by Eq. (9.298), because it indicates that, for the same value of c\ud835\udf032 = \ud835\udf092, \ud835\udf032 becomes positive if \ud835\udf0e2 = + 1 and \ud835\udf032 becomes negative if \ud835\udf0e2 = \u2212 1. The corresponding values of \ud835\udf031 are determined by Eqs. (9.301) and (9.302). (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the wrist joint variables (\ud835\udf035, \ud835\udf034, \ud835\udf036)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure1-1.png", + "caption": "Fig. 1. Parts of the Waveboard.", + "texts": [ + " Finally, the multibody model is enhanced by modeling the wheels as two tori, performing a comparison between the ring and toroidal models. \u00a9 2020 Elsevier Ltd. All rights reserved. The Waveboard, also called Casterboard, Ripstik or Essboard, is a variant of the skateboard with two platforms (known as decks) that are coupled by a torsion bar and allowed to rotate independently about the same axis. Each of these decks lies over a passive caster, enabling the relative motion between the wheels and the platforms ( Fig. 1 ). The Waveboard is a highly movable dynamical system, with five degrees of freedom, and an interesting propelling mechanism: forward motion is achieved by performing a lateral oscillatory motion, without touching the ground. Mechanics and control of robotics and biological locomotion have been widely studied by several authors [1,2] . In particular, a well-known mechanism in the context of nonholonomic systems similar to the Waveboard is the Snakeboard. It is characterized by a higher degree of simplicity and stability, being its analysis much simpler" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001769_robio.2015.7418980-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001769_robio.2015.7418980-Figure3-1.png", + "caption": "Fig. 3 structure of the adjusting mechanism", + "texts": [ + " According to the theorem that 3 points determine a plane, the equation of the pressure plate plane can be expressed as 0Ax By Cz D (2) As p1, p2 and p3 are all in the plane, we can obtain 1 1 1 0Ax By Cz D (3) 1 1 2 0Ax By Cz D (4) 1 1 3 0Ax By Cz D (5) Combine (3), (4) and (5) 1 3 3 2 1 2 2 2 2 z z D C z z A C d z z B C e (6) By Combining (2) and (6) , we can obtain the equation of the pressure plate plane. 2 3 2 1 1 3 1 1 2 0 z z z z x y z z z x y (7) Then we can obtain the normal n 2 3 2 1 1 1 , , ( , , 2) z z z z n A B C x y (8) Meanwhile, the equation of n can also be expressed as 0 0 0x y z A B C (9) Where 2 3 1 z x A z , 2 1 1 z y B z , C=2. B. Principle of drill attitude adjusting Adjusting attitude is the key technology in automatic drilling system[6][7]. As is shown in Fig.3, there is the principle of adjusting mechanism, which is composed of a spherical pair, a spherical pair bearing and two eccentric discs. The radii of two eccentric discs both are r, and the line O is the geometry axis of the big eccentric disc. While the line Ob is the eccentric axis of the big eccentric disc and it is also the geometry axis of the small eccentric disc. The eccentric axis of the small eccentric disc is the line Os. The drill centerline crosses the spherical pair and spherical plain bearing which is installed at the eccentric point of the small eccentric disc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002949_s12541-019-00277-9-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002949_s12541-019-00277-9-Figure8-1.png", + "caption": "Fig. 8 Anti-backlash dynamic model of the NTCIG", + "texts": [ + " 3 NLTE Calculation of\u00a0NTCIG with\u00a0Anti\u2011backlash Figure\u00a06 illustrates the anti-backlash principle of the NTCIG in that the meshing tooth thickness of the gear pair on every transverse plane changes when the driven gear moves along the axial direction. The relationship between the axial displacement (s) and the backlash (Bh) can be expressed as follows: (10)kb( ) = kd(\u2212 ) 1 3 Based on the displacement increment along the line of action shown in Fig.\u00a07, the detailed computations of NLTEd, NLTEb, and Bh for the NTCIG with eccentricity and center distance errors can be expressed as: [1, 21] 4 DTE Calculation of\u00a0NTCIG with\u00a0Anti\u2011backlash Figure\u00a08 illustrates the dynamic model of the NTCIG with a dynamic equation that is expressed as [23]: (12) NLTEdanti = NLTEd + s tan R 2 NLTEbanti = NLTEb + Bh \u2212 s tan R 2 (13) \ufffd \u0394F 1d \u0394F 2d \ufffd = \ufffd e1 0 0 e2 \ufffd\ufffd sin( 1 + + 10 ) \u2212 sin( + 10 ) sin( 2 \u2212 + 20 ) \u2212 sin(\u2212 + 20 ) \ufffd \ufffd \u0394F 1b \u0394F 2b \ufffd = \ufffd e1 0 0 e2 \ufffd\ufffd sin( 1 \u2212 + 10 ) \u2212 sin(\u2212 + 10 ) sin( 2 + + 20 ) \u2212 sin( + 20 ) \ufffd NLTEd = 180 \u2217 60 R 2 cos \u22c5 2\ufffd i=1 \u0394Fid NLTEb = 180 \u2217 60 R 2 cos \u22c5 2\ufffd i=1 \u0394Fib Bh = 180 \u00d7 2 sin R 2 cos \ufffd e 1 e 2 \u0394a \ufffd\u23a1\u23a2\u23a2\u23a2\u23a3 \u2212 cos( 1 + 10 ) cos( 2 + 20 ) 2 sin \u23a4 \u23a5\u23a5\u23a5\u23a6 In the anti-backlash process, the driven gear moves towards the dotted line position from the solid line position so that an additional angle is produced, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003847_0954405420978120-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003847_0954405420978120-Figure5-1.png", + "caption": "Figure 5. Adjusting vectors computation of VCWPIA and VCLSPIA: (a) VCWPIA method and (b) VCLSPIA method.", + "texts": [ + " In which, gk i = aki bk i , ak i =1 bk i gk i . P= P0 PN 2 4 3 5, Qk = Qk 0 Qk N 2 4 3 5, Ak = 1 0 0 0 . . . 0 bk 1 ak 1 gk 1 0 . . . 0 0 bk 2 ak 2 . . . . . . 0 .. . . . . . . . . . . . . . .. . bk N 1 ak N 1 gk N 1 0 0 0 . . . 0 1 0 BBBBBBB@ 1 CCCCCCCA , Dk = Dk 0 Dk N 2 4 3 5 \u00f019\u00de The first method is VCWPIA (Variable Coefficient Weighted Progressive Iterative Approximation).The adjusting vector can be represented by equation (20). The adjusting vector of VCWPIA considers the error vector of one target point, as shown in Figure 5(a). Dk =P Ak 1Qk 1 \u00f020\u00de The second method is VCLSPIA (Variable Coefficient Least-Squares Progressive Iterative Approximation). The adjusting vector of VCLSPIA considers the error vector of adjacent three target points, which can be represented by equation (21). The diagram can be seen in Figure 5(b). Dk =Ak 1T P Ak 1Qk 1 , \u00f021\u00de The range of step mk are computed according to the eigenvalue of Ak 1 or Ak 1TAk 1.9 For the two methods, the VCWPIA is prefer to be used as an interpolation method while VCLSPIA as an approximation method. An example is used to compare smoothing path generated by the two methods using same tolerance constraints. As shown in Figure 6, the input linear path has five target points. The position path using VCWPIA and VCLSPIA are illustrated in Figure 6(a), the rotation angle path relative to the parameter of position arc length are plotted in Figure 6(b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-FigureD.13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-FigureD.13-1.png", + "caption": "Fig. D.13 Elliptical plate as a particular case of a elliptical prism", + "texts": [ + " This fact is less evident when comparing the inertia moments in terms of the mass as in (D.38) and (D.24) because the the mass is already scaled, being the mass of the elliptical prism 8 times the mass of it corresponding one-octant segment. Example D.13 (The inertia tensor of a ellipsoidal plate) An ellipsoidal plate is a particular case of an elliptical prism for which the height of the prism is very small compared to the ellipsoidal face dimensions. Then its inertia moment matrix arise after (D.24) assuming a \u2248 0 (Fig.D.13): Example D.14 (The cylinder) The cylinder is also a particular case of the elliptical prism where both main diagonal components of the elliptical plane are equal. For instance b = c = r implies an horizontally placed cylinder of radius r and length l = A = 2a, with volume V = \u03c0r2l, and diameter D = B = C = 2r . The inertia matrix is found after these conditions over expression (D.24) (Fig.D.14). 430 Appendix D: Examples for the Center of Mass and Inertia Tensors of Basic Shapes I c = m 12 \u23a1 \u23a3 6r2 0 0 0 l2 + 3r2 0 0 0 l2 + 3r2 \u23a4 \u23a6 = m 48 \u23a1 \u23a3 6D2 0 0 0 4l2 + 3D2 0 0 0 4l2 + 3D2 \u23a4 \u23a6 (D" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001273_ccdc.2013.6561824-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001273_ccdc.2013.6561824-Figure4-1.png", + "caption": "Fig. 4. Structure of Falcon UUV Fig. 5. Thruster Distribution Total vector of propulsion forces and moments in the horizontal plane can be written as [10], [16]:", + "texts": [ + " Considering the difficulty of computing ce in (6), a feedback control input of acceleration error is introduced c c ck= \u2212e e (7) To eliminate chattering problem caused by the discontinuous term, an adaptive term is added in the control law to replace the switching term \u02c6 ( ) 2ad est K= + + \u039b \u03c4 \u03c4 C s (8) est\u03c4 is an adaptive term that estimates the lumped uncertainty vector \u03c4 . The estimation of the lumped uncertainty vector is proposed to follow: est = \u0393\u03c4 s (9) The total control law can be defined as \u02c6 ( ) 2eq ad eq est C K= + = + + + \u039b \u03c4 \u03c4 \u03c4 \u03c4 \u03c4 s (10) In this paper, our mainly work focus on the cascaded tracking control with fault thruster reallocation of FALCON UUV (from Lab. of Underwater Vehicles and Intelligent System). Fig. 4 shows its structure. A brief sketch of the vehicle's horizontal thruster distribution is shown in Fig. 5. This UUV has four horizontal thrusters, which are fixed in symmetric layout, denoted as , [1, 4]iHT i \u2208 , and one 4916 2013 25th Chinese Control and Decision Conference (CCDC) vertical thrusters. The thruster configuration of FALCON enables direct control of 4 degrees of freedom (DOF): surge, sway and yaw control in the horizontal plane and heave control in the vertical plane. As only one thruster in the vertical plane, there is no need for fault tolerant control in the vertical plane and horizontal thruster fault tolerant control is discussed in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002687_mwscas.2016.7869969-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002687_mwscas.2016.7869969-Figure3-1.png", + "caption": "Fig. 3. Coordinate systems and forces/moments acting on the quadcopter (m = 1.656 [kg], g = 9.80665 [m/s2], I11 = 0.01982 [kg \u00b7 m2], I22 = 0.01954 [kg \u00b7 m2], I33 = 0.03221 [kg \u00b7 m2], L = 0.365 [m])", + "texts": [ + " (20) Here, r and f denote the position vector on which external forces act and a vector of the external forces, respectively. The coordinate components of an angular momentum of the quadcopter is given as Eq. (21). HA = 3\u2211 i=1 HiAEi. (21) Then, we have the following equation: (H1,H2,H3)T = I\u0302(\u21261,\u21262,\u21263)T. (22) The moment of inertia I\u0302 is also defined by I\u0302 = I11 I12 I13 I21 I22 I23 I31 I32 I33 . (23) In this paper, for I\u0302, we assume that Ikl = 0, and for k , l. III. Dynamic model of a quadcopter The coordinate systems and free body diagram for the quadcopter are shown in Figure 3. Based on the preceding mathematical foundation, we can describe the dynamic model of the quadcopter (Fig. 3). Fi, (i = 1, 2, 3, 4) and Mi, (i = 1, 2, 3, 4) of Figure 3, represent vertical forces and moment, respectively. Fi, (i = 1, 2, 3, 4) and Mi, (i = 1, 2, 3, 4) are defined in a similar manner [1]. Each motor of the quadcopter has an angular speed \u03c9i and produces a vertical force Fi according to: Fi = kFi\u03c9 2 Mi, i = 1, 2, 3, 4, (24) Experimentation with a fixed motor in a steady state shows that kF \u2248 1.79\u00d7 10\u22127 N rpm2 . The motors also produce a moment according to: Mi = kMi\u03c9 2 Mi, i = 1, 2, 3, 4, (25) The constant, kM , is determined to be approximately 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001264_tmag.2013.2268737-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001264_tmag.2013.2268737-Figure1-1.png", + "caption": "Fig. 1. Conic closed surface symmetric about the axis with spherical surfaces", + "texts": [ + " For a wedge shaped closed surface, the total force from Maxwell\u2019s stress equation is null. The forces on each surface of the wedge closed surface provides some general insight into Maxwell\u2019s stress equation. Section II-C evaluates Maxwell\u2019s stress equation for two separated current elements. The forces on each individual current at and . element as a result of Maxwell\u2019s stress equation are not necessarily equal and opposite. Maxwell\u2019s stress (3) may be used to determine the force on each surface of the conic closed surface shown in Fig. 1. The surfaces at and are spherical surfaces (i.e., only a radial component normal to the surface) and the side surface only has an azimuthal component normal to the surface. The angle specifies the tilt of the side surface with respect to the axis. The force from Maxwell\u2019s stress equation for surface is (13) and the force in the z direction for surface is (14) The force from Maxwell\u2019s stress equation for surface is (15) and the force in the z direction for surface is (16) The force fromMaxwell\u2019s stress equation for surface is (17) and the force in the direction for surface is (18) For classical electrostatics, the total force from Maxwell\u2019s stress equation on a conic surface outside a differential charge element at the origin is null [8]", + " This nonzero result supports the reality that the Biot-Savart law \u201chas meaning only as one element of a sum over a continuous set of a current loop or circuit\u201d [9], yet the analysis in this section is for a single isolated differential current element. A variation of Maxwell\u2019s stress equation for magnetostatics based on one of the possible differential force elements of (10) is the most likely candidate for a meaningful stress equation involving a single differential current element. If such a variation of Maxwell\u2019s stress equation exists, it is presumed to yield a null force for any arbitrary enclosed surface (including the conic surface of Fig. 1) where the single differential current element is not on the surface. From a Cylindrically Symmetric Infinite Line Current on the Axis With Current, , in the Positive Direction Maxwell\u2019s stress equation for magnetostatics is only applicable for closed circuits. The simplest \u201cclosed circuit\u201d to evaluate is the infinite line current. For an infinite cylindrically symmetric line current on the axis with total current in the positive direction, the magnetic field external to the line current is (19) where is the distance from the axis", + " Equations (63) and (64) are now used to solve for the [8] that may be used to cancel Maxwell\u2019s stress equation for the spherical surface: (65) In a similar manner, a may be derived such that when the variant of Stoke\u2019s theorem (34) is applied yields a spherical surface stress of (66) where is is an unknown constant related to the current element. The magnetostatic recast stress equation of (66) manifests an inward surface tension only in the radial direction (referenced to the differential current element location) where the units of the constant are N/sr. The net force on any arbitrary closed surface in free space (including the conic surface of Fig. 1) is null. This is easily shown for a concentric spherical surface with the differential current element at the origin of the sphere due to symmetry. The net force on any arbitrary surface in free space that doesn\u2019t contain the differential current element can also easily be shown to be null by applying the variant of Stokes\u2019 Theorem and the method of prior work for an arbitrary surface [8]. The magnetostatic Maxwell stress (per unit area) of (50) at concentric spherical surfaces (with radius ) away from the differential current element falls off as ", + " Using a variant of Stokes\u2019 Theorem, Maxwell\u2019s stress equation for an isolated current element may be recast eliminating the tensile stress in the direction of the magnetic field and establishing a stress that is normal to the magnetic field, directed inward toward the differential current element: (69) where is an unknown constant related to the current element. The magnetostatic recast stress equation of (69) manifests an inward surface tension only in the radial direction (referenced to the differential current element location). The net force on any arbitrary closed surface in free space (including the conic surface of Fig. 1) is null. The magnetostatic Maxwell stress (per unit area) of (50) at concentric spherical surfaces (with radius ) away from the differential current element falls off as . In contrast, the magnetostatic recast stress equation of (69) falls off as . The radial stress of (69) may be modeled as an average rate of photon influx and out-flux to/from the differential current element accounting for the pinch stress identified in prior work [30]. For two separated differential current elements, an alternate magnetostatic stress equation is established from the historical force formulas for differential current elements and the results from prior work for electrostatics [30]: (70) The magnitude of the line stress is equivalent to the two current element Neumann force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003066_j.promfg.2020.04.133-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003066_j.promfg.2020.04.133-Figure6-1.png", + "caption": "Fig. 6. Basic sequence (a), case 1: RBPI without delay time (b), case 2: RBPI with delay time of 10s (c)", + "texts": [ + " Based on the analysis, the time delay, welding parameters and even the welding sequence can be adjusted. The decision matrix takes the beads, which are not deposited yet out of the sequence of the basic simulation into account and sorts them according to the temperature analysis. In this paper, the algorithm decides only for the parameter welding sequence, the others (time delay, welding parameters) are predefined and remain constant. Two cases for the RBPI are investigated. The first case is for a predefined delay time of 0s and the second of 10s (Fig. 6). Both cases are compared with the basic path sequence. One layer has a welding time of 9s and is partitioned into 5 beads with a length of 12mm. The welding time of one bead takes 1.8s. Fig. 7 shows the welding sequence of the first three layers of the RBPI algorithm. The first layer is predefined by the boundary conditions as explained earlier and the second layer starts at t = 9s. It can be seen, that the lowest temperature position is the starting point of the bead. Therefore, the algorithm positions the first two beads of the second layer in the same order as in the first layer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001805_s11771-015-2685-5-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001805_s11771-015-2685-5-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of experimental device: (a) Front view; (b) Perspective view", + "texts": [ + " On the other hand, DENG et al [28] proposed a three-dimensional finite element model of a theoretical assembling straight bevel gear pair in order to investigate the contact fatigue and bending fatigue at the level of tooth surface and tooth root respectively. We are interested in the failure of gear tooth under the effect of solid contaminant. In fact, the evolution of temperature, wear and vibration levels in gear mechanism is studied. On the other hand, we have tried to obtain a better understanding and a good description of debris distributions by using unimodal, single distribution models. An original experimental device (Fig. 1) was used to carry out the rolling\u2212sliding contact experiments. The device has been designed for the investigation of pinion/gear contacts. The pinion was actuated at a rotational speed n1=500 r/min by an electric motor. The gear was actuated in rotation under the driving pinion action and loaded by a friction torque. To apply the friction torque on the gear, a simple mechanism of braking was used. In fact, the principle consists of braking the rotational movement of the gear, with a clamping mechanism which presses a steel disc in contact with the gear, thus preventing its rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003119_s12206-020-0130-7-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003119_s12206-020-0130-7-Figure8-1.png", + "caption": "Fig. 8. Contact pressure varies with tilting angle, \u03b8t = 0.02\u02da: (a)-(d) Lundberg\u2019s logarithmic formula; (e)-(h) two-side tilt case.", + "texts": [ + " (21)), the logarithmic crowning model that takes into account of two-side tilt is obtained, and the symmetry of cylindrical structure is maintained. However, when the logarithmic formula of one-side tilt is superposed, both sides of the crowned roller contain the smaller crown drop of original relaxed section, and then the half of the superposed crown drop will inevitably reduce the crown drop of original squeezed section. In order to verify the effectiveness of this superposition-average weighting method, the contact characteristics of the cylindrical roller profiled in case of two-side tilt are discussed. The result is shown in Fig. 8, normal load q is 0.8P, and \u03b8a is 0\u02da, 0.01\u02da, 0.02\u02da and 0.03\u02da, respectively. In Fig. 8, the contact pressure of two crowned rollers varies regularly with the tilting angle \u03b8a. In the alignment or misalignment state, the effective contact length of the roller that crowned base on the angular misalignment is slightly shorter. For the roller with Lundberg\u2019s logarithmic profile, the stress edge effect becomes prominent with the increase of the normal load q and the tilting angle \u03b8a, which will result in a decrease of contact fatigue life of the bearing. when the tilting angle \u03b8a (\u03b8a = 0.03\u02da) exceeds the designed angle \u03b8t (\u03b8t = 0.02\u02da), as shown in Fig. 8, there is no obvious stress edge effect at ends of the roller profiled in case of two-side tilt, and the pressure varies uniformly and no distortion occurs, which indicates that the crowned roller still have well adaptability to the angular misalignment state. Referring to Figs. 8, 9(a) and (b), when the tilting angle \u03b8a = 0\u02da, the contact force distribution of two crowned rollers is symmetrical, while the contact pressure of the roller crowned under two-side tilt is gentler at roller ends. In Fig", + " 9(c), variations of the maximum contact pressure and deformation with the normal load are consistent with that in above analysis, but within entire normal load range, the contact pressure and deformation of the roller crowned in case of twoside tilt change a little compared with the profile in case of oneside tilt. Compared with the roller crowned with Lundberg\u2019s formula, the elastic deformation is increased by 21.65 %~8.60 (corresponding load q from small to large) when \u03b8a = \u03b8t = 0.001\u02da, and is increased by 1.47 %~4.37 % when \u03b8a = \u03b8t = 0.02\u02da. When the load q reached 1.1P, no significant stress edge effect occurred. This illustrates the effectiveness of the superposition-average weighting. Compared with Fig. 8, the logarithmic formula for two-side tilt case can satisfy the profile crowning of the rollers in heavy-duty cylindrical roller bearings, for it can avoid the stress edge effect and improve the pressure distribution under the large load and tilting angle. And it can also be used as a crowning method for rollers that do not need to consider the misalignment problem. The stress field of working surface and subsurface of the cylindrical contact pair has a direct impact on early peeling and service life of bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002003_cjme.2015.0713.092-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002003_cjme.2015.0713.092-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of a journal bearing", + "texts": [ + " The experimental results show that the shaft vibration characteristics are different from the vibration characteristics beside the journal in that process. A non-linear model of the fluid film force is used to numerically relate the rotational, whirling and squeezing speeds of the journal, and then the vibration properties of the journal. The experiments show the influences of the lubricating water flow velocity on the first critical speed\uff0c the time when significant oil-film whirling motion begins and the vibration amplitude at the bearing center. The motion of the journal in a sliding bearing is illustrated in Fig. 2. The journal center velocity can be resolved into the rotational speed, \u2126=d\u03c6/dt, the whirling speed, \u03c9=d/dt, and the radial squeezing velocity, \u03bd=d\u03b5/dt. The analytical function of the oil film pressure in the sliding bearing with the dynamic boundary condition[21] can be described by Eq. (1): { } 2 1 2 3 1 02 0 , 2 ip f f f z c z c r = + + + + (1) ( ) ( ) 1 2 sin sin2 3 1 cos 1 cos i i i i f = + - + + ( )2 3 sin2 , 1 cos i i + (2) ( ) ( )2 2 3 sin sin12 1 cos1 cos i i ii f = - - ++ ( ) ( ) 2 2 2 3 sin cos sin cos6 2 , 1 cos 1 cos i i i i i i + + + (3) ( ) ( )3 2 3 sin cos12 1 cos1 cos i i ii f = - + ++ ( ) ( ) 2 2 2 3 sin sin6 2 , 1 cos 1 cos i i i i - + + (4) where pi is the film pressure in the film\uff0c\u03bc is the kinetic viscosity of the fluid\uff0cr0 is the journal radius, \u03b5 is the journal center eccentricity, \u03bb is the ratio of the average clearance to the journal radius(c/r0), \u03c6i is the circumferential angle on the journal surface\uff0cz is the axial coordinate, c1 and c2 are the integration constants when using the bearing inlet and outlet conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003524_j.addma.2020.101612-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003524_j.addma.2020.101612-Figure18-1.png", + "caption": "Fig. 18. SEM fractographs: (a) SLM rod, (b) annealed SLM rod, and (c) commercial rod.", + "texts": [ + " Whereas, the commercially available 304 used in this study is mainly composed of FCC austenite (\u03b3-Fe) single phase that is typical of a wrought and annealed structure. As the ferrite/austenite phase boundaries in the SLM structure effectively hinder dislocation movement and increase the overall strength. The resistance of the material to indentation is a qualitative indication of the strength. As the hardness test includes a substantial component of plastic deformation, the hardness generally shows a strong correlation with tensile strength as shown in Table 4. The images presented in Fig. 18 show the SEM fractographs of tested rods displaying a dimpled structure. A large number of crater-like voids are observed on the fracture surface of the SLM rod, as shown in Fig. 18 (a). These voids are believed to be the critical sites for failure initiation during tensile testing, and responsible for the relatively lower ductility of the SLM rod. During the SLM process, relatively weak bonded regions may remain due to insufficient melting, although this was not detected by CT scanning. These weak bonded regions would be pulled out during tensile testing, leaving behind crater-like voids on the fracture surface. The number of such voids has been minimized by the annealing treatment, as shown in Fig. 18(b). Although the SLM rod may contain unavoidable weak regions, its mechanical properties can be improved by annealing. One feature observed in the fractographs is that the average size of the annealed SLM rod is much smaller than that obtained in the commercial rod shown in Fig. 18(c). In general, the coalescence of a set of micropores occurring on the site of defects results in larger dimples during a tensile test [60]. In the annealed SLM rod, the formation of T.W. Hwang et al. Additive Manufacturing 36 (2020) 101612 dimples is related to the plastic deformation at each micropore site without coalescence, and resulting in the small-sized dimples observed on the fracture surface. Herein, an innovative manufacturing process to fabricate fine metallic rods via selective laser melting (SLM) was developed and the following conclusions were obtained: 1) A forming process that comprised of SLM, laser rounding, and sizing was successfully implemented using AISI 304 stainless-steel powder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001169_gt2014-26891-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001169_gt2014-26891-Figure3-1.png", + "caption": "Fig. 3 Coupled rotors with angular misalignment", + "texts": [ + " The resulting equation of motion of each element is { } { } { } { } 11 15 1 1 1 51 66 1 11 15 1 1 1 51 66 1 11 66 11 15 51 11 0 00 0 0 0 000 0 0 0 00 0 00 0 0 0 0 00 0 0 0 00 0 00 0 00 , , e ye e e y e e e e e e m m M G q q m m M G k k k q Q k k k A EHere k k k k k k l \u23a1 \u23a4\u23a1 \u23a4 \u23a2 \u23a5\u23a2 \u23a5\u23a1 \u23a4 \u23a1 \u23a4\u23a3 \u23a6 \u23a2 \u23a5\u23a3 \u23a6\u23a2 \u23a5 + \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5\u23a1 \u23a4 \u23a1 \u23a4\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6\u23a3 \u23a6 \u23a3 \u23a6\u23a3 \u23a6 \u23a1 \u23a4 \u23a2 \u23a5\u23a1 \u23a4\u23a3 \u23a6\u23a2 \u23a5+ = \u23a2 \u23a5 \u23a2 \u23a5\u23a1 \u23a4\u23a2 \u23a5\u23a3 \u23a6\u23a3 \u23a6 = = = = && & (4) Incorporating proportional damping for shaft material and rewriting element eq. (4) in abbreviated form [ ]{ } [ ]{ } [ ]{ } { }e e e e e e eM q c q k q Q+ + =&& & (5) Assembling different matrices of beam elements of driving rotor, the resulting equation of motion is [ ]{ } [ ] { } [ ]{ } { }1 1 1 1 1 1 1M q C q K q Q+ + =&& & (6) The displacement vector corresponding to eqs. (4-6) is [ ]e x yq z x y\u03b8 \u03b8\u23a1 \u23a4= \u23a3 \u23a6 In case, the angular misalignment is present between the rotors as shown in Fig. 3, the coordinate systems attached with driver and driven rotor shafts do not have the same orientation. Therefore, coordinate transformation of second rotor is required before the assembly of mass, stiffness, damping matrices and force vector of two coupled rotors. Considering the coordinate system attached with the driver shaft parallel to global coordinate system, the nodal displacement vector of each element of the driven rotor is transformed using transformation matrix i.e. [L]. In the present work, the angular misalignment (\u03b8) has been considered only in XZ plane so the nodal transformation matrix is [ ] 0 0 0 0 0 0 0 0 1 0 0 0 0 0 1 0 0 0 0 0 1 l m m l L \u23a1 \u23a4 \u23a2 \u23a5\u2212\u23a2 \u23a5 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 (7) Here l = Cos \u03b8, and m = Sin \u03b8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000479_3dp.2015.0011-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000479_3dp.2015.0011-Figure1-1.png", + "caption": "Figure 1. Jaws filament drive gear and mid - section view of Jaws filament drive gear.", + "texts": [ + " We suspect the reliability of the filament feed gear is dependent upon three factors: (1) size of contact surface between the drive gear and the filament; (2) depth of the gear \u2019 s tooth engagement into the filament; (3) number of teeth engaged in the filament at any one time; and (4) the direction of the force vector imparted by the filament drive gear into the filament. R&D at re:3D has delved into this problem and below is the result of their work. The goal of this work is to characterize the amount of force that can be generated\u00a0 by a machined extruder drive\u00a0 bolt, affectionately named Jaws.\u00a0 Figure 1 shows a 3D rendering of\u00a0 the filament drive gear, Jaws, that will\u00a0be mounted to the shaft of a geared NEMA 17 stepper motor. The testing outlined below was performed with a Greg \u2019 s Wade \u2013 type extruder for 3\u00a0 mm polylactic acid (PLA) filament (see Fig.\u00a02). The Jaws filament drive gear is machined using a four - axis computerized numeric control milling machine. In the design process, our aim was to optimize the four variables stated above. \u2022 The amount of contact between the drive gear and the filament is maximized by machining the gear teeth using a cutting tool of the same diameter as the filament it will drive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001131_icrom.2013.6510116-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001131_icrom.2013.6510116-Figure3-1.png", + "caption": "Figure 3 - Virtual Double Inverted Pendulum", + "texts": [], + "surrounding_texts": [ + "The dynamics and kinematics model of a two wheel mobile manipulator robot with a reaction wheel is described in this section. For the considered robot which is shown in Figures (1-2), its parameters specifications are given in Table (1). The base of the mobile manipulator is assumed as a passive joint. Therefore, the manipulator part consists of three links and actuators in 3D workspace. The reaction wheel is placed on the first link which has one more degree of freedom rather than this one. The dynamics equations of motion of a mobile manipulator are described as: '( = M(q)q + H(q,q) + G(q) (1) where q=[ql'q2'\" .. ,qtER7 is the vector of generalized coordinates and =[\" 2,0, 4' 5' 6' lER7 is the input generalized torque and M(q) ER7x7, H(q, q) ER7xI, G( q) ER7X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively. The considered two WMM robot with a reaction wheel required complex dynamics modelling as a result of under actuated system. The passive joint causes that the balancing challenge of this robot in the XOY plane is more important than other its positions. Thus, Double Inverted Pendulum Model (DIPM) is utilized to simplify the dynamic analysis for the balancing control, [18]. The effectiveness of this simplified model is demonstrated for the dynamic locomotion of the highly nonlinear and complex system, [19]. Tablel - Parameters of the two WMM with a reaction wheel o XYZ OI_XIYIZI Q,&q2 Q3 Q;(i=4,5,6) q, L 1,(i= I ,2,3) 14 Word coordinate frame Mobile manipulator coordinate frame Rotation angles of wheels Inclination angle of the passive joint Joint angles of links Angular position of the reaction wheel Radius of wheels Distance between wheels Length of links Distance of the reaction wheel centre to first joint B. Dynamics Modelling of Double Inverted Pendulum It is possible to model two WMM as a virtual double inverted pendulum model. In this model, the components of second and third of manipulator in XOY plane are considered as the virtual second link of double inverted pendulum model. This simple model is used as the stabilizing control part. This model is shown in Figure (3) and parameters of these models is stated in Table (2). Table2 - Parameters of DIPM ilw Rotation angle of wheels ill Joint angle of the first link ile Joint angle of the second link il2 Angular position of the reaction wheel 11 Length of the first link Ie Length of the second I ink 12 Distance of reaction wheel centre to first joint xme X position of the me in 01 coordinate frame Yme Y position of the me in 01 coordinate frame Xl X component of the first link Length Y1 Y component of the first link Length me Equivalent mass on the second link m1 Mass of the first link m2 Mass of the reaction wheel mw Mass of wheels c,(i=2,e) Coefficient of friction a1 Length of the CoG on the first link i1 Moment inertia of link I around the ZI axis i2 Moment inertia of the reaction wheel around the zraxis Calculation of positions of the virtual link xmeand Yme' and its length Ie' joint angleqe and the mass meare as following: (2) (3) (4) (5) The dynamics equations of motion of this model are obtained by Euler-Lagrange equation as: (6) whereq = [qw, q1' q2' qeF ER4 is the vector of generalized coordinates and r = [fw, f1' f2' feF ER4 is the input generalized torque for the virtual double inverted pendulum modeI.M(q) ER4X4,A(q, q) ER4XI,G(q) ER4X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively which are calculated as: M = [\ufffd\ufffd\ufffd M31 M41 where: Mll = (mw + m1 + mz + me)rZM12 = M21 = (a1m1 + me (11 + Ie COS(qe)) + 11mz)rcOS(q1) + mwrZM14 = M41 = melerCOS(q1 + qe) Mzz = i1 + iz + a1 Z m1 + 11 z mz + 11 z me + Ie z me cosz qe + 211 leme cos qe + mWrZM23 = M32 = M33 = izMz4 = Ie z me cosz qe + 11 leme cos qe M13 = M31 = M34 , Z = M43 = OM44 = Ie meH1 = -fiI(mer(ft1 sin(i'iI) (\\1 + Ie COS(qe)) + leqe sin(cl!+qe)) + 11mZft1rsin(Q1) + a1 m1 ib r Sin(Q1)) - merqe(leqe Sin(q1 + qe) + leq1 COS(q1) sin(qe))Hz = -2m.leq1 (feqe COS(qe) sin(qe) + 11 qe sin(qe)) + meleqe (qwr COS(q1) COS(qe) - qe sin(qe) (\\1 + 21e cos(qe))) H3 = cZqZH4 = -me Ie rih qw COS(qe) sin(qe) + m.leq1 z sin(qe) (11 + Ie COS(qe)) + CeqeG1 = G3 = OGz = -a1 m1 Sin(q1) - 11 sin q1 (mz + me) - Ie me (cos q1 sin qe + sin q1 cos qe)G4 = -Ieme sin(q1 + qe) C. Dynamic Verification/or DIPM ADAMS is the most widely used multi-body kinematics and dynamics analysis software in the word. Also, Adams helps engineers to study the kinematics and dynamics of moving parts, how loads and forces are distributed throughout mechanical systems, and to improve and optimize the performance of their products. The extracted dynamics equations of motion of the virtual double inverted pendulum with a reaction wheel are verified by the ADAMS model which the 3D sketch of DIPM is shown in Figure (4).We insert some similar torque as input to these two models and compare the reaction of model joints with together. This torques is shown in Figure (5) which insert to the reaction wheel. Moreover, the obtained result of the model response is shown in Figure (6).These curves show that the dynamics equations of motion of the virtual double inverted pendulum are verified. So, we can use these equations in the control algorithm to realize the dynamic stability. In addition, the parameters specifications of this system in this verification routine are expressed in Table (3). 0.4 0.3 0.2 E 0.1 ;;. ! -0.1 -0.2 -0.3 I I I I I I I I I ___ 1 ____ 1___ -.l ___ .l ___ L ___ 1 __ _ 1___ -.J ___ \ufffd __ _ I I I I -- \ufffd --- 1 --- T --- \ufffd -- -0.4 - - -1- - - -1 - - - -+ - - - + - - - t- - - - 1- - - -1- - - ---j I I I I I I I I -0.5 OL--- :0c'c .5=- -L.- 1\ufffd.5=- -\ufffd2 --2cc'.C:-5-\ufffd3---=-3.L5 --':-\ufffd4\"'.5-\ufffd 5 Time (8) Since, the two WMM contains a passive joint, it requires an active controller to stabilize the passive joint and to control its stability. Related to ZMP criterion,[20], this robot have just one stable natural position that the moment of the gravity has no effect on the body, in other words, gravity force direction of COG crosses the mobile wheel axis. If we control the position of the COG to this point, we can able to satisfy the robot stability. In this section, a PID controller is proposed to control the motion of the reaction wheel to achieve dynamic stability. The supervising controller identifies the position of the COG and tunes the set point of PID controller. At the same time, the PID controller moves the reaction wheel to reach the stable natural position. Also, the control block diagram of this controller is shown in Figure (7). The supervising controller finds the COG position without any information about robot links and the related body. So, as shown in Figure (7), the input parameters of this block are the reaction wheel angular velocity and its acceleration. Therefore, it can just find the position of the COG to the right or the left from the normal position of the first link (qi=O) with this information. For an example, in Figure (2), if the reaction wheel angular velocity and its acceleration are positive, it is clear that the position of COG are in the left side from the normal position of the first link (qi=O).SO, the supervising controller changes the PID set point to the negative value. Thus, this can use many type of controller for this block such as PI, PID, Fuzzy or GA controller. In this paper we used the PI controller for this block. IV. SIMULATION RESULTS AND DISCUSSION The validation of the proposed control strategy is demonstrated by simulation results. This simulation runs at MA TLAB Simulink Toolbox for the assumed robotic system which the parameters specifications are expressed in Table(3). In addition, the controller gains are expressed in Table (4). Table4--Controller gains amount Kp Coefficient of P action 3.916 KI Coefficient of I action 1 KD Coefficient of 0 action 0.0624 Ksd Coefficient of q3 parameter -6e-3 Ksp Coefficient of q3 parameter 0.0209 This type of robot is dynamically stable and this stability is attained using a reaction wheel. Here, we consider the three bench marks and run the simulations. In the first and second one, the robot start from an initial position and make itself stable and, in third one, the robot run in the different initial position and pass the variable acceleration. The initial conditions of these case studies are respectively expressed in Table (5). Moreover, the accelerate variation curve of the third case is shown on Figure (8). Table 5 - Initial conditions for simulations Casel Casell CaselII q qp i'i2 Joint angle of the first link (deg) -30 -30 30 Joint angle of the second link (deg) -60 -120 90 Angular position of the reaction wheel (deg) 0 0 0 1.5 - - - - - I - -----1- - - - - - 1 - - --- 1 ----- - - - - - ---1 - -----1- - - - - - r- - ---- _____ --J ______ 1 ______ L ____ _ ........ 0.5 1 .\ufffd \ufffd ] -0.5 I I I ----- T -----1-----l------------ \ufffd -----r----- \ufffd----- \ufffd------------ \ufffd.5 -----\ufffd----- 4----- \ufffd------------ -20\ufffd------\ufffd,0\ufffd----\ufffd2\ufffd0 ----\ufffd3\ufffd 0------\ufffd4 0\ufffd----\ufffd 5\ufffd 0------\ufffd6 0 Time(s) Figure 8 - Acceleration variation during the operation As shown from simulation results in Figures (9-17), the time history of the robot motion is presented in Figures (9- 11) for the cases to move and reach to the stable position. Moreover, the position of the <11 must be varied until the CoG reaches to the appropriate position Figures (12-14). In this position, the robot is stable and there is no moment appears to change this style. Also, the PID modifies the set point value to satisfy the stabilization. Figures (15-17) are shown the position of the reaction wheel, during the robot reaches to stable position. These results show that the reaction wheel movement makes the internal moment that it can control the passive joint. In this control strategy, the balancing is not achieved by the robot movement and only the reaction wheel is used to stabilize. This method can also improve the dexterity of the robot motion and increases the robot controllability." + ] + }, + { + "image_filename": "designv11_34_0000440_icuas.2015.7152421-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000440_icuas.2015.7152421-Figure1-1.png", + "caption": "Fig. 1. Acceleration Mapping", + "texts": [ + " Appropriate states for a given trajectory: x\u0304r = [ xir yir zir \u02d9xir \u02d9yir \u02d9zir ]T (8) Actual states of the pendulum tip: x\u0304 = [ xi yi zi x\u0307i y\u0307i z\u0307i ]T (9) State space for computing the LQR gain matrix: \u02d9\u0304x = Ax\u0304+Bu\u0304 (10) A = [ 03\u00d73 I3\u00d73 \u00b7 \u00b7 \u00b7 03\u00d76 ] and B = [ 03\u00d73 I3\u00d73 ] Desired acceleration given the LQR gain matrix K: u\u0304 = \u2212K(x\u0304r \u2212 x\u0304) (11) Input for a given trajectory: u\u0304r = [ x\u0308ir y\u0308ir z\u0308ir ]T (12) upm = u\u0304+ u\u0304r (13) Total input to the point mass system: ud = upm + [ 0 0 g ]T (14) To realize the desired acceleration, upm, on the pendulum tip we require a mapping from input variables to system state variables. This is provided given the assumption that the quadrotor can only act on the pendulum tip in the direction along the pendulum shaft as seen in figure 1. Thus, the required system states can be found as follows: xvr = L ud(1) \u2016ud\u2016 (15) yvr = L ud(2) \u2016ud\u2016 (16) pnr = xi \u2212 xvr (17) per = yi \u2212 yvr (18) hr = zi \u2212 \u221a L2 \u2212 x2vr \u2212 y2vr (19) \u03c8r = 0 (20) The following have been set to zero since an appropriate mapping would require estimating a desired jerk applied to the pendulum tip. \u02d9xvr = \u02d9yvr = ur = vr = wr = 0 (21) The second tier in our hierarchal controller seeks to control the state of combined system of the inverted pendulum and quadrotor. We use the same methods as those described in the first tier" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002222_ciec.2016.7513774-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002222_ciec.2016.7513774-Figure1-1.png", + "caption": "Figure 1: Model of TRMS", + "texts": [ + " The results show that the proposed design strategies are competitive to the conventional design techniques. The rest of the paper is organized as follows. First the TRMS is introduced and mathematical model is presented in section-II. Section-III introduces the brief description of PSO, a stochastic search algorithm and then the state feedback controller and observer is designed for TRMS system. 978-1-5090-0035-7/16/$31.00\u00a92016IEEE 30 Section-IV presents the simulation results and section-V concludes the paper. II. TWIN ROTOR MIMO SYSTEM The Two Rotor MIMO System (TRMS), as shown in Fig. 1, is a laboratory set-up designed for control experiments. From the control point of view, it exemplifies a high order nonlinear system with significant cross couplings. Twin rotor MIMO system has its two units such as mechanical unit and electrical unit. Two rotors attached to a balancing beam with a counter balance attached with pivot mounted on a tower make the mechanical unit. Two dc motors with two tachometers to sense the speed are positioned on these two rotors and one position sensor is mounted on the pivot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003708_bia50171.2020.9244518-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003708_bia50171.2020.9244518-Figure3-1.png", + "caption": "Fig. 3. (a) Mesh quality and dimensions the distal shaft of the catheter, (b) deformed shape of the catheter subjected to a 5 N tip force obtained from FE analysis.", + "texts": [ + " (12) This inverse solution reveals the external force fx and fy through knowledge of the deformation (image-processing) and mechanical properties of the catheter. The flexural rigidity of the catheter was set to EI = 723.7 MPa.mm2 [5]. III. VALIDATION STUDY In order to validate the proposed image-based method, the force estimation was performed on the results of a series of finite element (FE) simulations. The image-processing and solution were implemented in Matlab 2019b. The FE simulations were performed in Ansys v17 (Ansys Inc., PA, USA). Fig. 3(a) shows the FE model solved in this study. The model was solved with nonlinear geometry (large deformation) assumption and had 51 nodes and 50 linear bending elements. Homogeneous Dirichlet and Neumann boundary conditions were applied at the left-most node of the model to simulate the cantilever condition. A vertical concentrated force of 0.5\u20135 N magnitude was applied to the right-most node in separate simulations. Fig. 3(b) shows the representative deformed shape of the catheter under 5 N. B. Image-based Deformation Extraction To extract the shape of catheter, first each image was converted to gray-scale and thresholded manually to obtain a binary map of the pixels belonging to the catheter. Afterward, the centerline of the extracted shape (1-pixel width) was obtained through averaging the vertical white pixels, i.e., pixels with \u20191\u2019 value. Fig. 4(a) shows the binary map and the extracted centerline of Fig. 3(b). The parameter s was defined for each individual pixel on the centerline as: si = 1 L \u2211i k=1 \u221a (xk \u2212 xk+1)2 + (yk \u2212 yk+1)2 i = 1 \u00b7 \u00b7 \u00b7 p\u2212 1, (13) where, p was the total number of pixels constructing the centerline, x and y were the horizontal and vertical position of the pixel in the image in millimeters. Next, to obtain coefficients of the shape interpolation functions introduced in Eq. 1, an iterative curve-fitting was performed on < si, xi > and < si, yi > ordered 2-tuples. The curve-fitting was stopped once the goodness-of-fit would surpass a minimum R2 \u2265 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003302_i2mtc43012.2020.9129036-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003302_i2mtc43012.2020.9129036-Figure5-1.png", + "caption": "Fig. 5. The test rig.", + "texts": [ + " Fast Fourier transform (FFT) is used to obtain the frequency spectrum. The frequency spectrums of blade displacement signal are shown in figure 3. And the frequency spectrum of the BSC signal is shown in figure 4. It can be seen that the frequency spectrum of BSC contains the frequency components of two blade displacement signals. In this section, we will verify the effectiveness of the proposed model on experimental data. NLSF is used to process the BSC signal collected by non-uniformly spaced sensors and generate frequency spectrum. The test rig is shown in figure 5. Assume the angle interval between TOA sensor a and TOA sensor b is \u2206ab. Three TOA sensors are used, and \u220612 = \u220623 = 30\u25e6, as shown in figure 6. The rotating speed is 5280 RPM. Notice that the angle intervals of the sensors have nothing to do with whether the BSC can be used for BTT. Different BTT signal processing methods may require specific sensor layouts, such as the autoregressive method [3] requiring four sensors arranged at equal intervals. The blade displacement signals calculated by using OPR as a reference are shown in figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003248_s00170-020-05597-z-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003248_s00170-020-05597-z-Figure8-1.png", + "caption": "Fig. 8 Verification experiment of guide joint", + "texts": [ + " The stiffness value obtained by the THK Company is an average stiffness (Fa/\u03b4a), and the stiffness value predicted by the present model is an exact stiffness obtained by derivation (\u2202Fa/\u2202\u03b4a). Therefore, the predicted value of stiffness is obtained when the axial load is an average load (1.5 times the preload). The approximation will cause some errors, but it is acceptable. The values predicted by the present model are in good agreement with the experimental values of THK Company, and some comparison results are listed in Table 5. The verification experiment of the guide joint is shown in Fig. 8. The modal test is shown in Fig. 8a. The frequency response functions obtained by different excitation positions in the modal test are shown in Fig. 8b. According to the calculation parameters in Table 2, the predicted normal and tangential stiffness are 3.02 \u00d7 109 N/m. The finite element model is shown in Fig. 8c, and the normal and tangential stiffness of the guide joint are entered into the COMBIN 14 element. The comparison between theoretical and experimental values for the guide joint is shown in Table 6. Sun et al. [25] identified the contact stiffness of the guide (SRG65LV2SSC0) based on experimental modal analysis, and their identification result is 3.18 \u00d7 109 N/m, which is also in good agreement with the prediction result of the present model. The verification experiment of the fixed joint is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002961_0036850419897221-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002961_0036850419897221-Figure14-1.png", + "caption": "Figure 14. Simplified model diagram of porous media.", + "texts": [ + "20 Porosity after deformation: e= 1 d2n D0 D1\u00f0 \u00de 2 D0 2 D1 2 B cos u \u00f07\u00de Viscous resistance coefficient: an = 80 1 e\u00f0 \u00de2 e3d2 \u00f08\u00de The viscous resistance coefficient and inertia resistance coefficient before and after the brush deformation are substituted into the steady state fluid analysis model of porous media model. Because these two coefficients defined in fluent are expressed by space vector and the whole brush ring is symmetrical in the center axis, the brush ring at the center angle of the 1/20 at 18 is taken in the calculation, as shown in Figure 14. Then, the leakage of the brush seal is simulated again with the porous medium model. The results are shown in Table 1. Table 1 shows he calculation results under different pressure differences. The second column is the average deformation of the filament after the fluid\u2013solid coupling calculation. The third column is the porosity value calculated after entering formulas (7). Finally, the table shows the calculation results about the leakage without deformation and the leakage with deformation and their differences" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000846_j.ymssp.2014.11.015-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000846_j.ymssp.2014.11.015-Figure2-1.png", + "caption": "Fig. 2. Free-body diagram of the follower; refer to Fig. 1 for the two coordinate systems.", + "texts": [ + " (6) and (7) with respect to time, _\u03c8 i\u00f0t\u00de and _\u03c8 j\u00f0t\u00de are obtained as follows: _\u03c8 i\u00f0t\u00de \u00bc \u03c70\u00fee cos \u03b1\u00f0t\u00de \u03b10 _\u03b1\u00f0t\u00de r \u0394o\u00f0t\u00de \u00fe0:5wb\u00fe2rd sin \u03b1\u00f0t\u00de \u03b10 _\u03b1\u00f0t\u00de \u00fer \u0394o\u00f0t\u00de cos \u03b1\u00f0t\u00de\u00fe\u03c6\u00f0t\u00de _\u03b1\u00f0t\u00de\u00fe _\u03c6\u00f0t\u00de \u00fe _r \u0394o\u00f0t\u00de sin \u03b1\u00f0t\u00de\u00fe\u03c6\u00f0t\u00de e cos \u03b1\u00f0t\u00de\u00fe\u0398\u00f0t\u00de _\u03b1\u00f0t\u00de\u00fe _\u0398\u00f0t\u00de ; \u00f08\u00de _\u03c8 j\u00f0t\u00de \u00bc \u03c70\u00fee sin \u03b1\u00f0t\u00de \u03b10 _\u03b1\u00f0t\u00de\u00fe r \u0394o\u00f0t\u00de \u00fe0:5wb\u00fe2rd cos \u03b1\u00f0t\u00de \u03b10 _\u03b1\u00f0t\u00de \u00fer \u0394o\u00f0t\u00de sin \u03c6\u00f0t\u00de\u00fe\u03b1\u00f0t\u00de _\u03c6\u00f0t\u00de\u00fe _\u03b1\u00f0t\u00de _r \u0394o\u00f0t\u00de cos \u03c6\u00f0t\u00de\u00fe\u03b1\u00f0t\u00de e sin \u03b1\u00f0t\u00de\u00fe\u0398\u00f0t\u00de _\u03b1\u00f0t\u00de\u00fe _\u0398\u00f0t\u00de : \u00f09\u00de Here, _r \u0394o\u00f0t\u00de \u00bc 0:5ab b2 a2 sin 2\u0394o\u00f0t\u00de _\u03c6\u00f0t\u00de _\u0398\u00f0t\u00de h i a sin \u0394o\u00f0t\u00de 2\u00fe b cos \u0394o\u00f0t\u00de 2h i1:5 : \u00f010\u00de The angle \u03c6(t) corresponding to the contact point Oc is determined at every instant for a given \u03b1(t) and \u0398(t) by locating the point on the elliptic profile of the cam which is tangential to the follower. Hence the slope of the follower, sb(t)\u00bctan( \u03b1(t)), should be equal to the slope of the cam at Oc (scO\u00f0t\u00de) which is calculated as follows: scO\u00f0t\u00de \u00bc tan \u0398\u00f0t\u00de\u00fe tan 1 b2 a2 tan \u03c6\u00f0t\u00de \u0398\u00f0t\u00de \" #\" # \u00f011\u00de Equating sb(t) and scO\u00f0t\u00de and rearranging, \u03c6(t) is calculated by the following: \u03c6\u00f0t\u00de \u00bc\u0398\u00f0t\u00de tan 1 b2 a2 tan \u03b1\u00f0t\u00de \u0398\u00f0t\u00de \" # \u00f012\u00de The equation of motion of the follower when it is in contact with the cam is derived by balancing the moments (from Fig. 2) about P as IPb \u20ac\u03b1\u00f0t\u00de \u00bcmbglg cos \u03b1\u00f0t\u00de\u00f0 \u00de Fs\u00f0t\u00dedx\u00feFn\u00f0t\u00de\u03c7\u00f0t\u00de Ff \u00f0t\u00de 0:5wb\u00fe2rd\u00f0 \u00de: \u00f013\u00de Here, IPb is the moment of inertia of the follower about P, mb is the mass of the follower, g is the acceleration due to gravity, and \u03c7(t) is the moment arm of the contact force about the pivot P. The elastic force from the spring, Fs(t), is given by the following, where Lus is the original length of the follower spring: Fs\u00f0t\u00de \u00bc ks Lus dy\u00fedx tan \u03b1\u00f0t\u00de\u00f0 \u00de\u00fe0:5wbsec \u03b1\u00f0t\u00de\u00f0 \u00de : \u00f014\u00de The normal force (Fn(t)) arising from the point contact with the cam is given by Fn\u00f0t\u00de \u00bc k\u03bb \u03c8 i\u00f0t\u00de \u03c8 i\u00f0t\u00de c\u03bb _\u03c8 i\u00f0t\u00de: \u00f015\u00de The non-linear contact stiffness is defined for a point contact based on the Hertzian contact theory [23] as k\u03bb \u03c8 i\u00f0t\u00de \u00bc \u00f04=3\u00deYe \u03c1e\u00f0t\u00de \u03c8 i\u00f0t\u00de 0:5 : \u00f016\u00de Here, Y is Young's modulus (with superscript e denoting equivalent) in accordance with the Hertzian contact theory given by the following, where \u03bd is Poisson's ratio, Ye \u00bc 1 \u03bd2c Yc \u00fe1 \u03bd2b Yb \" # 1 : \u00f017\u00de The equivalent radius of curvature at the contact (\u03c1e(t)) and the radius of curvature of the elliptical cam at Oc (\u03c1c(\u0394o(t))) are given by \u03c1e\u00f0t\u00de \u00bc \u03c1c \u0394o\u00f0t\u00de 1\u00fe\u00f0rd\u00de 1 h i 1 ; \u00f018\u00de \u03c1c \u03940\u00f0t\u00de \u00bc a sin \u03b3 \u0394o\u00f0t\u00de 2\u00fe b cos \u03b3 \u0394o\u00f0t\u00de 2h i1:5 ab : \u00f019\u00de The contact damping is modeled as linear viscous damping as the system is designed to not undergo any impact", + " The \u03bc is estimated from measured reaction forces (Nx(t) and Ny(t)) along the e\u0302x and e\u0302ydirections, respectively, and the tangential acceleration (lb \u20ac\u03b1\u00f0t\u00de) of the follower at its free end. By dividing the measured tangential acceleration by lb, \u20ac\u03b1\u00f0t\u00de is obtained, and then numerically integrating it twice w.r.t. time, the time-varying component of \u03b1(t) is computed, while the integration constant (\u03b1d) is obtained from the time-averaged value of \u03b1k(t). The instantaneous elastic force Fs(t) is calculated from \u03b1(t) using Eq. (14). From Fig. 2, Nx(t) and Ny(t) are evaluated as follows: Nx\u00f0t\u00de \u00bc Fn\u00f0t\u00de sin \u03b1\u00f0t\u00de\u00f0 \u00de\u00feFf \u00f0t\u00de cos \u03b1\u00f0t\u00de\u00f0 \u00de mblg \u20ac\u03b1\u00f0t\u00de sin \u03b1\u00f0t\u00de\u00f0 \u00de; \u00f027\u00de Ny\u00f0t\u00de \u00bc Fn\u00f0t\u00de cos \u03b1\u00f0t\u00de\u00f0 \u00de Ff \u00f0t\u00de sin \u03b1\u00f0t\u00de\u00f0 \u00de\u00fembg mblg \u20ac\u03b1\u00f0t\u00de cos \u03b1\u00f0t\u00de\u00f0 \u00de Fs\u00f0t\u00de: \u00f028\u00de Rearrange Eqs. (27) and (28) to yield the friction and normal forces as Ff \u00f0t\u00de \u00bcNx\u00f0t\u00de cos \u03b1\u00f0t\u00de\u00f0 \u00de\u00fe mbg Ny\u00f0t\u00de Fs\u00f0t\u00de sin \u03b1\u00f0t\u00de\u00f0 \u00de; \u00f029\u00de Fn\u00f0t\u00de \u00bcmblg \u20ac\u03b1\u00f0t\u00de\u00feNx\u00f0t\u00de sin \u03b1\u00f0t\u00de\u00f0 \u00de\u00fe Ny\u00f0t\u00de\u00feFs\u00f0t\u00de mbg cos \u03b1\u00f0t\u00de\u00f0 \u00de: \u00f030\u00de Since the dynamic force transducer used does not measure the DC component, a technique to estimate \u03bc is proposed that utilizes complex-valued Fourier amplitudes while maintaining the phase relationship among the measured signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003026_j.icte.2020.03.007-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003026_j.icte.2020.03.007-Figure4-1.png", + "caption": "Fig. 4. Mini maps showing position of drones in experiment.. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", + "texts": [ + " When an evader drone performs a nonattacking movement such as hiding, this movement will not create any damage to the secured area (num dmg). We can neglect this movement according to our objective. Because our main objective is on decreasing num dmg, we focus our experiment on showing that our proposed algorithm can perform a good result in handling a kamikaze drone direct attack. This kind of approach is computationally formal in promoting an optimization algorithm for MPME problem, as used on [5]. In a kamikaze drone maneuver, each evader drone movement is intended to make a direct attack towards a target. Fig. 4(a) shows the initial position of evaders and pursuers in each testing. We set the nearest evader distance from the secured area to 40 units. Please note that Fig. 4 is just a mini-map from the sky-view, which symbolizes every drone with a circle dot. If we take a closer look at the simulation environment, then Fig. 5 shows how we manifest the drones. In Fig. 4, evaders are symbolized with the brown dots, while pursuers are symbolized with the green dots. The red circle in the center of Fig. 4 indicates the central point of the area needed to be secured. Evaders have two main formations from four directions, as shown in Fig. 4(a), which are triangle formation as shown on the left and right side of Fig. 4(a), and rectangle formation as shown on the upside and downside of Fig. 4(a). Each evader starts from different (random) height, then moving directly but in a constant height to a secured area. Every evader does not perform any hiding or dodging maneuver because it will create a bias on performance evaluation. Here, we follow the evaders\u2019 movement equation as used on [5]. When an evader is near enough to its target, it changes its latitude to the height of the secured area. Meanwhile, the pursuers start from the random position from surrounding of a secured area. After some period of iteration, the coordinate of each drone may change, as shown in Fig. 4(b). Please notice that the bigger the size a dot has in Fig. 4, the higher its position from the ground. Fig. 5 shows that there are two kinds of drones in the simulation environment. In this experiment, the green drones represent pursuer drones; meanwhile, the brown drones represent evader drones. The purple line animates the process of a pursuer capturing an evader. If we take a closer look at Fig. 5, it tends to be complicated for pursuer drones to handle all of the evaders if the pursuers do not have a good strategy to solve the problem. Thus, to efficiently manage the MPME issue discussed here, a switching target communication strategy is proposed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003320_j.cmpb.2020.105646-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003320_j.cmpb.2020.105646-Figure12-1.png", + "caption": "Fig. 12. The corrected trajectory of the incisal pin.", + "texts": [ + " Indeed, measurement always involves error and there is always tradeoff between increasing the accuracy and reducing the cost. Since we need to position the marker outside of the mouth, the two markers are positioned in front of the upper jaw and lower jaw respectively as closely as possible. This will not create large lateral protrusion errors because the lateral movement usually involves smaller angles (Bennet angles < 20 \u00b0) but may create larger errors in the z direction because of larger opening angles ( > 40 \u00b0) and larger z direction movement in protrusion (see Fig. 12 ). The enlargement of error in the z direction is similar to the Abbe error (magnification of angular error over distance) in industrial metrology cases. Fortunately, for our dental application, we are not dealing with accuracy of micron level but rather millimeter level. Therefore, the result is near sub-millimeter range and is considered acceptable and we do not seek to use very high evel accuracy measurement or tracking device for cost reduction eason. .2. Movement simulation We develop the digital articulator system based on C ++ and penGL library", + " opening, protrusion, and lateral excursion. These movements re important for static and dynamic occlusal analysis for proshetic and orthodontic applications. All the movements are contrained by the condylar guidance to make sure the movement is onsistent with the real articulator. The system can load the digital teeth models acquired by 3Dcanning of the dental casts into the digital articulator system to imulate jaw movement. We also implement collision detection unction between opposing teeth. Fig. 12 shows the trajectory of he incisal pin moving from the origin to the protrusion position. here is an obvious ascent which responds to the contact of the pposing teeth, and this can provide us a predicted movement of ow the teeth contact in the patient\u2019s mouth. If there were no col- ision detection, the upper teeth would have gone through the op- osing teeth and caused wrong movement in the simulation. 6 m m v m c t p t t p r s l d i s D i w A N D S f R [ [ [ [ [ [ [ . Conclusions We develop a digital articulator which can simulate jaw moveent between opposing teeth and develop an optical tracking ethod to verify the accuracy of the digital articulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003439_s10854-020-04194-w-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003439_s10854-020-04194-w-Figure6-1.png", + "caption": "Fig. 6 Magnetic flux density distribution of transformer", + "texts": [ + " Secondly, the calculation results are loaded into the solid mechanics module and the overall force of the transformer is obtained by applying the boundary conditions. That whether the difference between the mechanical results calculated twice satisfies the iteration accuracy should be determined; if not, a recalculation should be made. Thirdly, we can judge whether the transient calculation time is complete or not. If not, another 1 time will be added and the above steps should be repeated. FE calculation results of magnetic distribution by COMSOL as shown in Fig.\u00a06 (instant time is 0.01\u00a0s). From Fig.\u00a06, at 0.01\u00a0s in a period, it is a special status while the magnetic field and magnetic flow are in symmetrical distribution between left and right limbs. Since the magnetic resistance of core made of silicon steel is far less than magnetic resistance of fixed structure and winding, the magnetic fluxes are concentrated mainly in the core. The measured 1 3 magnetization curves of silicon steel are used and the magnetostriction effect in the transformer numerical model is included by utilizing the magnetostriction curves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000668_fuzz-ieee.2013.6622512-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000668_fuzz-ieee.2013.6622512-Figure3-1.png", + "caption": "Fig. 3. TORA system", + "texts": [], + "surrounding_texts": [ + "We apply two methods in the previous section to two nonlinear systems." + ] + }, + { + "image_filename": "designv11_34_0003739_0954405420971128-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003739_0954405420971128-Figure7-1.png", + "caption": "Figure 7. Finite element models of the coin blanks and dies for the (a) \u2018Rooster\u2019 and (b) \u2018Eagle\u2019 prototype coins.", + "texts": [ + " For this purpose, all the contact interfaces of the coin blanks are defined by means of Nc pairs extracted from the faces of the three-dimensional elements that were used in their discretisation. The symbols gcn and gct denote the normal and tangential gap velocities in these contact pairs, which are penalised by large numbers K1 and K2 to avoid penetration. The above described procedure allowed modelling the coin blanks as deformable objects and required discretisation by means of non-structured three-dimensional meshes of hexahedral elements (Figure 7). The meshes of the coin blanks consisted of approximately 110.000 hexahedral elements with four layers of elements across thickness and a higher density of elements in the outer region where lettering was imparted. The dies and collar were modelled as rigid objects and were discretised by means of spatial triangular contact-friction elements. The meshes at centre of the dies are slightly recessed, as in the actual dies, so that contact with the coin blanks preferentially takes place at the outer region where lettering and additional design features are located (Figure 5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003225_tcyb.2020.2998505-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003225_tcyb.2020.2998505-Figure4-1.png", + "caption": "Fig. 4. TRAS system.", + "texts": [ + " (18) It is expected from the OEANFIS model to generalize learned knowledge well and to properly control a system in the real-time environment, by generating an adequate output signal x(t). The last unexplained element from Fig. 3 is the limiter, which is used to limit the input signal x(t) to a defined range. Its purpose is to protect a system from the presence of the possible unwanted input signal values. The limiter converts the signal from x(t) to x\u0303(t) and introduces it to the input of the system. The OEI controller proposed in the previous section will be experimentally tested on TRAS, manufactured by the company Inteco [44]. This laboratory model (Fig. 4) is a nonlinear multiple-input\u2013multiple output (MIMO) system and it represents a real challenge for stable control and optimization tasks. Communication between the system and a PC is provided by the RT-DAC/PCI I/O board. Control of the system is performed by using the MATLAB software package and Simulink toolbox. Authorized licensed use limited to: University of Exeter. Downloaded on June 19,2020 at 09:33:36 UTC from IEEE Xplore. Restrictions apply. The central part of the TRAS system is a metal beam that could be pivoted by a user (Fig. 4). The beam can rotate freely in horizontal and vertical planes, to achieve a specified position in 3-D space. Another vital part of the system is a counter-weight arm (counterbalance), which is fixed to the beam at the pivoting point. All rotations of the beam are provided by two dc motors, which are installed at both ends of it. The first motor is in charge of horizontal movements (it drives a tail rotor of the system) and the second motor controls movements in the vertical plane (and represents the main rotor of the system)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure5.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure5.3-1.png", + "caption": "Figure 5.3 A serial manipulator with an RRP arm.", + "texts": [ + " r\u0308o,a (o) + C\u0302(o,a)r\u0308a,b (a) + C\u0302(o,b)r\u0308b,c (b) + \u00b7 \u00b7 \u00b7 + C\u0302(o,p)r\u0308p,q (p) + 2[?\u0303?(o) o,aC\u0302(o,a)r\u0307a,b (a) + ?\u0303?(o) o,bC\u0302(o,b)r\u0307b,c (b) + \u00b7 \u00b7 \u00b7 + ?\u0303?(o) o,pC\u0302(o,p)r\u0307p,q (p)] + [\ud835\udefc(o)o,a + ?\u0303? (o)2 o,a ]C\u0302(o,a)r(a)a,b + \u00b7 \u00b7 \u00b7 + [\ud835\udefc(o)o,p + ?\u0303? (o)2 o,p ]C\u0302(o,p)r(p)p,q = r\u0308o,z (o) + C\u0302(o,z)r\u0308z,y (z) + C\u0302(o,y)r\u0308y,x (y) + \u00b7 \u00b7 \u00b7 + C\u0302(o,r)r\u0308r,q (r) + 2[?\u0303?(o) o,zC\u0302(o,z)r\u0307z,y (z) + ?\u0303?(o) o,yC\u0302(o,y)r\u0307y,x (y) + \u00b7 \u00b7 \u00b7 + ?\u0303?(o) o,rC\u0302(o,r)r\u0307r,q (r)] + [\ud835\udefc(o)o,z + ?\u0303? (o)2 o,z ]C\u0302(o,z)r(z)z,y + \u00b7 \u00b7 \u00b7 + [\ud835\udefc(o)o,r + ?\u0303? (o)2 o,r ]C\u0302(o,r)r(r)r,q (5.69) Figure 5.3 shows the side and top views of a serial manipulator, which is a typical example of an open kinematic chain. The arm of the manipulator consists of the first three links that extend up to the wrist point R. The end-effector of the manipulator is a gripper and it is attached to the arm at the wrist point with three successive revolute joints in a spherical arrangement so that their axes are concurrent at the wrist point. This example is concerned with the kinematic analysis of the arm only, excluding the end-effector", + " Complete kinematic analyses of serial manipulators with all the links will be considered later in Chapters 7\u20139. The links of the arm are connected with two successive revolute (R) joints and one prismatic (P) joint. The axes of the arm joints are represented by the following unit vectors. n\u20d701 = n\u20d710 = u\u20d7(0) 3 = u\u20d7(1) 3 , n\u20d712 = n\u20d721 = u\u20d7(1) 2 = u\u20d7(2) 2 , n\u20d723 = n\u20d732 = u\u20d7(2) 3 = u\u20d7(3) 3 The following reference frames are attached to the base and the moving links. Their basis vectors are indicated in Figure 5.3. 0(O),1(O),2(S),3(R) As noticed, the origins of the link frames are O0 = O1 = O,O2 = S,O3 = R For this system, the joint frames are assigned as indicated below. 01(O01) = 0(O),10(O10) = 1(O) 12(O12) = 1(S) \u2225 1(O),21(O21) = 2(S) 23(O23) = 2(S),32(O32) = 3(R) \u2225 2(S) The base frame origin O is selected on the axis of the first joint (01) at the level of the shoulder point S in order to minimize the number of the nonzero constant parameters. Thus, the only nonzero constant parameter happens to be the shoulder offset defined below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001862_icrom.2014.6990963-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001862_icrom.2014.6990963-Figure6-1.png", + "caption": "Fig. 6. The model of biped robot in Working Model 4D.", + "texts": [], + "surrounding_texts": [ + "In this section the pattern generation using GCIPM considering moving ZMP is proposed. Only the robot locomotion in single support phase (SSP) is considered. Also, it is supposed that the ZMP is moving only in x-direction during stepping and it is fixed in y-direction. So the ZMP trajectory in x-direction will be a time-variant function. Here, linear trajectory is considered for ZMP as (8).\n0 2( ) (0 ) 2 b Tx t t t T T \u239b \u239e= \u2212 \u2264 \u2264\u239c \u239f \u239d \u23a0\n(8)\nWhere, T is period of SSP and b is the range of ZMP movement under the supporting foot. It should be noted that the ZMP may move on the line connecting the heel to the toes without covering this line completely [10]. So b can be smaller than foot size. The trajectory generated by (8) for period of 4s is shown in Fig. 4. In this trajectory, b is 0.04 m. The trajectory generated by (8) is not the most natural ZMP trajectory, but it is similar to ZMP trajectory in human walking and is more natural than that of fixed ZMP. The ZMP trajectory in ydirection is the same as Fig. 2(b), because the ZMP is considered fixed in y-direction. Considering the ZMP trajectory as (8), the solution to (2) and (3) is:\n1 2 cost t fX A e A e t t\u03c9 \u03c9 \u03b2 \u03c1 \u03c9 \u03b1\u2212= + + + \u2212 (9)\n1 2 0 t tY B e B e y\u03c9 \u03c9\u2212= + + (10)\nWhere,\n( )1 1 1(0) (0) 2 A X X \u03b2 \u03b1 \u03c1 \u03c9 \u239b \u239e= + \u2212 + \u2212\u239c \u239f \u239d \u23a0\n(11)\n( )2 1 1(0) (0) 2 A X X \u03b2 \u03b1 \u03c1 \u03c9 \u239b \u239e= \u2212 \u2212 + \u2212\u239c \u239f \u239d \u23a0\n(12)\n1 1 (0)(0) 2 YB Y \u03c9 \u239b \u239e = +\u239c \u239f \u239d \u23a0 (13)\n2 1 (0)(0) 2 YB Y \u03c9 \u239b \u239e = \u2212\u239c \u239f \u239d \u23a0 (14)\n( )2\n2 2,\nb g\nT\n\u03c9 \u03bb \u03b1 \u03b2 \u03b1\n\u03c9\n+ = = (15)\n( )2\n0 2 2 f\nf\nL g z\u03bb \u03c9 \u03c1\n\u03c9 \u03c9\n\u2212 =\n+ (16)\n( ) 2 1 (2 4 )(0) (0) 1 1\nT T\nT T e T eX X e e\n\u03c9 \u03c9 \u03c9 \u03c9 \u03b2 \u03b1 \u03b2\u03c9 \u03b1 \u03c1 \u03c9 + \u2212= + \u2212 + + \u2212 \u2212 (17)\n1(0) (0) 1\nT\nT eY Y e\n\u03c9\n\u03c9\u03c9 += \u2212\n(18)\nIV. SIMULATIONS In this section, COM trajectories generated by GCIPM for fixed and moving ZMP are shown and compared with each other. The trajectory of biped joints for both fixed and moving ZMP are obtained solving inverse kinematics problem. Fig .5 shows a 6 degree of freedom (DOF) bipedal robot. This robot is used in simulations to solve the inverse kinematics and obtaining the joint trajectories during walking. It\u2019s assumed that the robots swinging foot is parallel with the ground surface during its walking. Also it is supposed that the upper body is upright during walking. Also, the simulations are done using Working Model 4D for fixed and moving ZMP.\nFig .6 shows the biped structure simulated in this software. It\u2019s supposed that the ZMP moves only in x-direction so the COM references in y-direction will be the same for both fixed and moving ZMP. That\u2019s why in this section the simulations are done only in x-direction to compare the trajectories for", + "fixed and moving ZMP. Simulations are done for T =1s, hf =0.1 m, L=0.2 m, b=0.04 m, M=20 kg, Hz =0.55 m and m=1 kg.\nFig. 7, show the ZMP and COM trajectories generated for both fixed and moving ZMP. COM trajectories generated by moving and fixed ZMP are illustrated in Fig. 8. It is seen that the trajectory generated considering the moving ZMP is smoother than fixed ZMP and it is near to straight line. Also considering b=0.04 m it can be conclude that the ZMP located under the supporting foot.\nFig. 9(a) and (b) represents the COM trajectories in different masses for fixed and moving ZMP, respectively. Comparing these figures, it can be concluded that the COM has less variation and stable motion in moving ZMP. For more comparison, the mean square errors (MSE) for different masses with fixed and moving ZMP are shown in Table. 1. It can be seen that MSE for moving ZMP is less than fixed one. To compare the joint trajectories for fixed and moving ZMP COM reference, the inverse kinematics problem is solved for 6-DOF bipedal robot shown in Fig. 5.\nThe joint trajectories are obtained as shown in Fig. 10. Simulations are done for li = 0.3m (i = 1,2,\u20265) and lab = 0.1m. As illustrated in Fig. 10, the joint trajectories corresponding to moving ZMP have the same pattern with fixed ZMP but with less variation in joints. To compare the torque needed in joints for walking in both fixed and moving ZMP cases, the results obtained using Working Model 4D are shown in Fig. 11. It is seen from Fig .11 that the joint torques in moving ZMP is decreased. So it can be stated that, the energy consumption for biped locomotion with trajectories generated for moving ZMP is less than the one generated for fixed ZMP and it has optimal gaits.", + "(a)\n(b)\nV. CONCLUSION In this paper, reference trajectory generation for biped, using GCIPM with moving ZMP has been proposed. By choosing a linear time-variant function instead of fixed point for ZMP reference, more natural trajectory for ZMP was considered. Applying this ZMP trajectory to GCIPM, the COM trajectory was obtained. The comparison of moving and fixed ZMP references indicated that for moving ZMP, COM trajectory is smoother. Also, solving inverse kinematics, the joint trajectories for 6-DOF bipedal robot were calculated. Results showed that trajectories for moving ZMP have the same pattern as fixed ZMP with less variation in joints. Biped" + ] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure20-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure20-1.png", + "caption": "Figure 20: Unpressurized nonrotating rotor interference CAE analysis with C3D8I elements for test seals \u2013 Displacement profile at 0.6mm rotor interference", + "texts": [ + " The rotor movement has been defined by applying enforced displacement to circular rotor\u2019s reference point. Rotor rub has been followed by pull back of the rotor to its original position to simulate loading and unloading steps conditions. 7 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use The displacement contour in the unpressurized-nonrotating rotor interference simulation with C3D8I elements is detailed in Figure 20 for 0.6 mm radial interference. Due to frictional effects and a cant angle of 45o, the rotor interference deflects the bristles both in the radial and tangential directions. The magnitude of the maximum bristle tip displacement is around 0.832 mm for a rotor interference of 0.6 mm. Usage of C3D8I elements instead of B31 elements enables better viewing of displacement profile with smoother contour. The VM stress profile at 0.6 mm rotor interference is given in Figure 21 for FE model using C3D8I elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000588_j.jfranklin.2015.03.036-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000588_j.jfranklin.2015.03.036-Figure1-1.png", + "caption": "Fig. 1. D region. (a) D region for vertical strips. (b) D region for disks.", + "texts": [ + " The normalized membership function is \u03bc \u03b2ij j xj t\u00f0 \u00de \u00bc w \u03b2ij j \u00f0xj\u00f0t\u00de\u00dePrj \u03b2ij \u00bc 1 w \u03b2ij j \u00f0xj\u00f0t\u00de\u00de \u00f03\u00de and 0r\u03bc \u03b2ij j \u00f0xj\u00f0t\u00de\u00der1; Xrj \u03b2ij \u00bc 1 \u03bc \u03b2ij j \u00f0xj\u00f0t\u00de\u00de \u00bc 1 \u00f04\u00de Then, the membership function \u03b1i\u00f0x\u00f0t\u00de\u00de of the ith fuzzy rule becomes \u03b1i\u00f0x\u00f0t\u00de\u00de \u00bc \u220f n j \u00bc 1 \u03bc \u03b2ij j \u00f0xj\u00f0t\u00de\u00de \u00f05\u00de and 0r\u03b1i\u00f0x\u00f0t\u00de\u00der1; Xr i \u00bc 1 \u03b1i\u00f0x\u00f0t\u00de\u00de \u00bc 1: \u00f06\u00de Using a center average defuzzifier, product inference, and incorporating fuzzy blending, the fuzzy system (2) can be described as follows: E _x\u00f0t\u00de \u00bc A\u00f0\u03b1\u00dex\u00f0t\u00de \u00fe B\u00f0\u03b1\u00deu\u00f0t\u00de \u00f07\u00de where A\u00f0\u03b1\u00de \u00bc Xr i \u00bc 1 \u03b1i\u00f0x\u00deAi; B\u00f0\u03b1\u00de \u00bc Xr i \u00bc 1 \u03b1i\u00f0x\u00deBi \u00f08\u00de Before giving our main results, some definitions and lemmas are given as follows: Definition 1. The fuzzy descriptor system E _x\u00f0t\u00de \u00bc A\u00f0\u03b1\u00dex\u00f0t\u00de \u00f09\u00de is regular, if det\u00f0sE A\u00f0\u03b1\u00de\u00dea0, 8 tZ0. Definition 2. The fuzzy descriptor system (9) is impulse-free, if det\u00f0sE A\u00f0\u03b1\u00de\u00de \u00bc rank E, 8 tZ0. The LMI region we considered in this paper is described as follows [35]: fD\u00bc zAC;R11 \u00fe R12z\u00fe RT 12zo0g \u00f010\u00de where R11 \u00bc RT 11ARd d and R12ARd d, d is called the order of the region D. Two typical regions are shown in Fig. 1 In Fig. 1(a), the corresponding characteristic functions can be described as fDa \u00bc 2a\u00fe z\u00fe zo0 when a\u00bc0, it becomes the left-half side of complex plane C . In Fig. 1(b), the corresponding characteristic functions can be described as fD\u00f0a;r\u00de \u00bc r a\u00fe z a\u00fe z r \" # \u00bc r a a r \u00fe 0 1 0 0 z\u00fe 0 1 0 0 T zo0 when a\u00bc0 and r\u00bc1, it becomes the disk centered with the origin at \u00f00; 0\u00de and radius 1. Definition 3. Given a region D, the fuzzy descriptor system (9) is D stable, if it is regular, impulse-free and there exists a Lyapunov function V\u00f0x\u00f0t\u00de\u00de satisfying R11 1\u00fe R12 1 2 _V \u00f0x\u00de V\u00f0x\u00de \u00fe RT 12 1 2 _V \u00f0x\u00de V\u00f0x\u00deo0 \u00f011\u00de Remark 1. This definition is adopted from [40]. The transient dynamic performances of the fuzzy system can be guaranteed by constraining the poles of the system in a prescribed region", + " The fuzzy descriptor system (14) is said to be stabilizable if there exist matrices Qi, Y, U, U1, M, Xi and a scalar \u03bc such that \u03a5 ii \u00bc \u03a3i 11 \u03a3i 12 n \u03bcUT \u03bcU \" # o Id I 0 0 0 ; i\u00bc 1; 2;\u2026; r \u00f043\u00de \u03a5 ij \u00bc \u03a3ij 11 \u03a3ij 12 n 2\u03bcUT 2\u03bcU \" # o 2 \u00f0r 1\u00de Id I 0 0 0 ; j\u00bc 1; 2;\u2026; r; ioj \u00f044\u00de where \u03a3 i 11 \u00bcUAT i \u00fe XT i B T i \u00fe AiU T \u00fe BiXi \u03a3 i 12 \u00bc \u00f0Qi \u00fe U1YU T 1 \u00de \u00fe \u03bc\u00f0AiU T \u00fe BiXi\u00de U \u03a3 ij 11 \u00bcUAT i \u00fe XT j B T i \u00fe AiU T \u00fe BiXj \u00fe UAT j \u00fe XT i B T j \u00fe AjU T \u00fe BjXi \u03a3 ij 12 \u00bc \u00f0Qi \u00fe U1YU T 1 \u00de \u00fe \u00f0Qj \u00fe U1YU T 1 \u00de \u00fe \u03bc\u00f0AiU T \u00fe BiXj\u00de \u00fe \u03bc\u00f0AjU T \u00fe BjXi\u00de 2U U1ARn \u00f0n p\u00de, satisfying ETU1 \u00bc 0, and Qi is defined as Eq. (31). Example 1. Consider the following fuzzy descriptor system: R1 : If x1\u00f0t\u00de is M1 1; Then E _x\u00f0t\u00de \u00bc A1x\u00fe B1u\u00f0t\u00de; R2 : If x2\u00f0t\u00de is M2 1; Then E _x\u00f0t\u00de \u00bc A2x\u00fe B2u\u00f0t\u00de; where E\u00bc 1 1 0 2 5 0 2 3:5 0 2 64 3 75; A1 \u00bc 4 1 0 g5 8 1 1 5 1 2 64 3 75; A2 \u00bc 4 1 1 4 8 3 1 5 1 2 64 3 75 B1 \u00bc 1 0:2 1 T ; B2 \u00bc 1 0:6 2 T and the region considered is as Fig. 1(a) with a\u00bc0.5. Using Theorem 2 and choosing \u03bc\u00bc 0:45, U1 \u00bc 10 11 14\u00bd T , a feasible solution is given as Q1 \u00bc 519:9254 331:2463 572:4514 331:2463 443:8743 489:4875 572:4514 489:4875 778:9983 2 64 3 75; Q2 \u00bc 519:9254 331:2463 572:4514 331:2463 333:7588 489:4875 572:4514 489:4875 888:7429 2 64 3 75; U \u00bc 113:8842 7:9850 75:7640 21:6782 16:1812 30:0014 104:7244 46:5637 192:8450 2 64 3 75; M \u00bc 51:2479 3:5933 34:0938 9:7552 7:2816 13:5007 47:1260 20:9537 86:7802 2 64 3 75; Y \u00bc \u00bd 3:4445 and Pi's as P1 \u00bc 175:4769 47:6471 90:2235 47:6471 27:0915 40:9632 90:2235 40:9632 103:8792 2 64 3 75; P2 \u00bc 175:4769 47:6471 90:2235 47:6471 83:0239 40:9632 90:2235 40:9632 213:6238 2 64 3 75; the consequent state feedback gains are as follows: K1 \u00bc \u00bd 8:5415 2:7644 9:0762 ; K2 \u00bc \u00bd 7:5421 6:1982 8:4471 Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001942_j.protcy.2016.03.017-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001942_j.protcy.2016.03.017-Figure1-1.png", + "caption": "Fig. 1. Configuration of the journal bearing showing the whirling orbit of the journal centre", + "texts": [ + "Re ijS Stiffness coefficients of micropolar fluid film, ,ri and ,rj , N/m ijS Stiffness damping coefficients of micropolar fluid film, LRCSS ijij 332 , ,ri and ,rj t Time, s U Velocity of journal, RU , m/s W Load in bearing, N 0W Steady state load in bearing, N 0W Non-dimensional steady state load in bearing, LRCWW 322 00 z Cartesian coordinate axis along the bearing axis, m z Non-dimensional Cartesian coordinate axis along the bearing axis, Lzz 2 Eccentricity ratio 0 Steady-state eccentricity ratio 1 Perturbed eccentricity ratio Whirl ratio, p Attitude angle, rad 0 Steady state attitude angle, rad zx, Micropolar fluid functions along circumferential and axial directions z, Non-dimensional micropolar fluid functions along circumferential and axial directions Characteristics length of the micropolar fluid Newtonian viscosity coefficient, Pa s p Angular velocity of the orbital motion of the journal centre, rad/s Angular velocity of journal, rad/s Circumferential coordinate, rad, ;Rx c Circumferential coordinate where the film cavitates, rad Non-dimensional time, t. 2. Analysis: 2.1. Modified Reynolds equation A schematic diagram of a hydrodynamic journal bearing with the circumferential coordinate system used in the analysis is shown in Fig. 1. The modified Reynolds equation in non-dimensional form [12-13] applicable for twodimensional flow of micro-polar lubricant with the turbulent effect under the dynamic condition is represented as follows: hh z p Nlh zL Dp Nlh mzm 21 2 1 ,,,, 2 (1) where, ; 2 coth 2 1 ,, 2 2 , 3 , Nlh l Nh l h k h Nlh m mmz mz z D zz B hCkhAk Re12 ,Re12 and The values of the turbulent coefficients, z CBA ,, and z D are obtained from references [6, 13]. It is assumed that the journal undergoes a whirling motion in an elliptical orbit about its mean steady-state position ( 00, ) with amplitudes ie1alRe and ie10alRe along the line of centres and perpendicular to the line of centres respectively", + " Stiffness and Damping coefficients With the dynamic pressure fields known, the non-dimensional components of stiffness and damping coefficients are obtained as follows: 1 0 0 1 cosalRe c zddpS rr ; 1 0 0 1 sinalRe c zddpS r (8a) 1 0 0 1 cosaginaryIm c zddpDrr ; 1 0 0 1 sinaginaryIm c zddpD r (8b) where, LRCSS ijij 332 and LRCDD ijij 332 The other four stiffness and damping coefficients rr DDSS and , , can be represented by analogy. 2.3. Stability characteristics The stability of the journal is analysed by combining the equations of motion and the resultant film forces rF and F in r and directions respectively. Referring to Fig. 1, the equation of motion of the rigid journal, assuming the rotor to be rigid, can be written as 0cos 2 2 2 0 dt d dt d MCWFr ; 0..2sin 2 2 0 dt d dt d dt d MCWF (9) For steady-state condition the equation of motion is written as 0cos 000 WFr ; 0sin 000 WF (10) Using the equations (2), (9) and (10), the following equations result in 00 0 0 20 sincos 1 rrrrrrrrrrr rr DD W DSSDDSSD DD WM (11) 0sincos cos 00 0 0 2 0 00 0 42 0 rrrrrrr rrrrrr SS W SSSS DDDDSS W WMWM (12) From the known values of stiffness and damping coefficients, the values of critical mass parameter and whirl ratio can be obtained by solving the above equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003272_j.promfg.2020.05.027-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003272_j.promfg.2020.05.027-Figure8-1.png", + "caption": "Figure 8. Maximum stress analysis at the corresponding resonant frequency.", + "texts": [ + " The vibration displacement vectors represent the exciting energy propagating alone the six waveguides to the front end with directional changes. At the front end of the vibrator, a steadystate hybrid L&T vibration can be observed (see Figure 7). To validate the mechanical property of the vibrator, this paper also analyzes the maximum stress of the vibrator at the resonant frequency under different exciting voltages. The study would be important for dynamic applications, especially on fatigue behavior. In the FE analysis, the maximum principal stress method was used to analyze the stress distribution, as shown in Figure 8. The maximum stress increases as the applied driving voltage are increased, as shown in Figure 9. The maximum voltage is 60 Vrms based on the capacity of PZT stacks and the power. According to the reported stress-life (SN) behavior of the SLM-AM AlSi10Mg parts, the part material can withstand 108 and more cycles under the stress 60MPa [16]. After heat treatment, the SLM-AM AlSi10Mg can sustain longer repetitive cycles under higher stress [15]. In our study, the predicted stress is far less than that in fatigue testing, and consequently, the theoretic fatigue life should be much longer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure14-1.png", + "caption": "Fig. 14. The surfaces where the HTCs are calculated, surface between the chosen inner air point and three internal surfaces.", + "texts": [ + " It is here estimated as, HTC = qs A (Ts \u2212 Tref ) (1) 870 Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 15,2021 at 16:48:44 UTC from IEEE Xplore. Restrictions apply. where qs is the heat flux, A is the surface area, Ts is the surface temperature and Tref is the reference temperature. The choice of reference temperature is not obvious. A point, P, in the air between the rotor and casing is chosen for the calculation of Tref , see Fig. 13. The surfaces where the HTCs are calculated are shown in Fig. 14. The results, seen in Table II, were used as an input to the LPN-model. TABLE II HTC VALUES OF SURFACES BETWEEN AIR AND END-WINDING, CASING AND ROTOR AS SHOWN IN FIG.14 Speed HTCair-ew HTCair-casing HTCair-rotor (rpm) (W/m2K) (W/m2K) (W/m2K) 1000 32 12 38 3000 62 35 67 5000 83 37 86 The variation of average component temperatures depending on outer boundary conditions is investigated for the full model of the spiral duct design. The outer BCs are varied between 0 to 45 W/m2K for one operating point (3000 rpm, 101 Nm). As can be seen in Fig. 15 the parts closest to the outer boundaries are the ones that are mostly affected (magnet, rotor core and shaft) with a \u2206T of 5-6oC whereas the winding temperatures do not change much (\u2206T of 2- 3oC and the stator core temperature is nearly constant (\u2206T is 1oC)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003860_j.euromechsol.2020.104197-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003860_j.euromechsol.2020.104197-Figure3-1.png", + "caption": "Fig. 3. The contours of out-of-plane displacement for the models with different moduli for the fibers in bottom layer at the final configuration: (a) \u03b3b = 1.0, (b) \u03b3b = 0.5 and (c) \u03b3b = 0.0.", + "texts": [ + " The results predicted by 800 solid-shell elements are almost consistent with ones by 3200 solid-shell elements, which means that refining models further almost do not affect the results and the displacement results converge for the models meshed with 800 solid-shell elements. Therefore, the computational costs and efficiency for the prediction of anisotropic swelling of hydrogel bilayers with different fiber moduli can be greatly decreased and improved by the solid-shell elements method, respectively. The contours of out-of-plane displacement for the models with different moduli of reinforced fibers in bottom layer at the final configuration are shown in Fig. 3. For the model with modulus of \u03b3b = 1.0, the final configuration with negative Gaussian curvature is similar with saddle surface. However, the square bilayer plates with fiber moduli of \u03b3b = 0.50 and 0.00 both curl upward. Therefore, the Gaussian J. Wang et al. European Journal of Mechanics / A Solids 86 (2021) 104197 J. Wang et al. European Journal of Mechanics / A Solids 86 (2021) 104197 curvatures of these two final configurations are positive. Meanwhile, the degree of helix and curling for the model with \u03b3b = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001297_j.jmatprotec.2013.01.009-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001297_j.jmatprotec.2013.01.009-Figure2-1.png", + "caption": "Fig. 2. Numerical", + "texts": [ + " The aim of this study was to develop a weld heat source applied to GTAW simulation which would enable the thermal field to be predicted and the fusion zone dimensions to be verified. It consisted in defining a simple analytical heat source, described by as few input parameters as possible: links between the numerical parameters of the model and the welding parameters were identified using two inverse methods. 2.2. Heat source definition In the GTAW process, the heat input is spread over a surface. This is the reason why a surface heat source described by homogeneous thermal flux Q was chosen: the welding power P was imposed uniformly through a radius disk R (Fig. 2). The thermal flux was given by the following relation (4): Q = P (4) \u00b7 R2 To perform a representative simulation, from a thermo mechanical point of view, it is necessary to describe the thermal field in the part during the welding process in a reliable way. In the next heat source. eratur p l s 2 w i F 2 p a t b t s b i t i m t a t T b the fusion zone dimensions. The correlation obviously required the aragraph, the welding parameters influencing temperature evoution in the part during the process are listed, thus the most ignificant ones could be identified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001576_ls.1326-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001576_ls.1326-Figure3-1.png", + "caption": "Figure 3. Geometry model of a micro-grooved bearing pad.", + "texts": [ + "19 The converging condition is confirmed by the following equation: P 1\u00f0 \u00de i;j Pi;j \u226410 10; i \u00bc 1; 2; ::; nk; j \u00bc 1; 2;\u2026;mk (1:7) where nk and mk are the numbers of nodes on the fourth layer along the \u03c6 direction and \u03bb direction respectively, Pi, j is the pressure value at the node (i, j) on the forth layer after a W-cycle is achieved and P 1\u00f0 \u00de i;j is the next pressure value of Newton\u2013Raphson iteration. Gas film thickness of the bearing pad The geometry model of a micro-grooved bearing pad is shown in Figure 3, where xs is the coordinate of the leading edge of the micro-grooved bearing pad along the x direction, lx is the length of the microgrooved bearing pad along the circumferential direction, \u03b3 is the orientation angle of the grooves, wg is the groove width, sg is the groove distance, hg is the groove depth and lg is the distance between two adjacent grooves. The dimensionless variables are defined in the following form. H0 \u00bc h0=c \u00bc 1\u00fe \u03b5 cos \u03c6; Hg \u00bc hg=c; Wg \u00bc wg=R; Lg \u00bc lg=R; Sg \u00bc sg=R; L1 \u00bc lx=R; L2 \u00bc B=R; \u03c6s \u00bc xs=R; (1:8) where h0 is the gas film thickness of the bearing pad without parabolic grooves, \u03b5 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2j \u00fe y2j q =c is the eccentric ratio, and xj and yj are the coordinates of the journal centre Oj along the x and y directions respectively (Figure 1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure4.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure4.2-1.png", + "caption": "Fig. 4.2 Roll, pitch and yaw angles [image obtained for internet]", + "texts": [ + " 3.9 The pair of rotation \u00f0\u201a; #\u00de: representing a rotation angle # about an arbitrary axis \u201a . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 161 Fig. 3.10 Scheme of the trigonometric functions of an angle h in a unit circle, Wikipedia (2013) . . . . . . . . . . . . . . . . . . . . . . . . . . . 165 Fig. 3.11 Rigid motion transformation . . . . . . . . . . . . . . . . . . . . . . . . . . . 165 Fig. 4.1 # rotation about an arbitrary axis \u201a\u00f0a; b\u00de . . . . . . . . . . . . . . . . . . 187 Fig. 4.2 Roll, pitch and yaw angles [image obtained for internet] . . . . . . 193 Fig. 4.3 Geometric interpretation for some Euler angles. Left: zxz convention. Right: zyz convention. First an a-rotation around z, so that the neutral axis N (either the new x or y axes) is properly aligned. Second a b-rotation around the N axis. Finally a -rotation around the newest z-axis so that the frame reach the final orientation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198 xx List of Figures Fig. 5", + "2 Roll-Pitch-Yaw Representation This is a non-symmetric (Tait\u2013Bryant parametrization), based-frame-defined (extrinsic) Euler angles representation, defined as the three successive rotations in the strict specific order x \u2212 y \u2212 z. The first rotation, performed about the xA-axis (the x-axis of the base frame) is called the roll angle, and is denoted by \u03c6. The second rotation, performed about the yA-axis (the y-axis of the base frame) is called the pitch angle, and is denoted by \u03b8. The third and last rotation, performed about the zA-axis (the z-axis of the base frame) is called the yaw angle, and is denoted by \u03c8 (Fig. 4.2). The attitudeparameter vector representing the three roll-pitch-yaw angles (RPY)of an arbitrarily moving frame 1 with respect to base frame 0 is defined as: \u03b8rpy \u239b \u239d \u03c6 \u03b8 \u03c8 \u239e \u23a0 \u2208 S \u00d7 S \u00d7 S = T 3 Since the basic rotations are all performed relative to the fixed frame 0 the composed rotation matrix R1 0 is defined as: R1 0(\u03b8) Rz,\u03c8Ry,\u03b8Rx,\u03c6 (4.12) Then replacing the basic rotations (3.16), (3.17) and (3.18), it yields to the particular solution of the Rotation Matrix using roll-pitch-yaw angles: R1 0(\u03b8) = \u23a1 \u23a3 c\u03c8 \u2212s\u03c8 0 s\u03c8 c\u03c8 0 0 0 1 \u23a4 \u23a6 \u23a1 \u23a3 c\u03b8 0 s\u03b8 0 1 0 \u2212s\u03b8 0 c\u03b8 \u23a4 \u23a6 \u23a1 \u23a3 1 0 0 0 c\u03c6 \u2212s\u03c6 0 s\u03c6 c\u03c6 \u23a4 \u23a6 = \u23a1 \u23a3 c\u03c8c\u03b8 \u2212s\u03c8c\u03c6 + c\u03c8s\u03b8s\u03c6 s\u03c8s\u03c6 + c\u03c8s\u03b8c\u03c6 s\u03c8c\u03b8 c\u03c8c\u03c6 + s\u03c8s\u03b8s\u03c6 \u2212c\u03c8s\u03c6 + s\u03c8s\u03b8c\u03c6 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8c\u03c6 \u23a4 \u23a6 (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000995_s12206-013-0808-1-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000995_s12206-013-0808-1-Figure7-1.png", + "caption": "Fig. 7. An impeller part model for the case study.", + "texts": [ + " In fact, a DMT can be modeled as a 3-D geometry that consists of multiple 2-D sliced layers accumulated, each layer of which is also built by bonding multiple adjoined clad geometries with each other according to 2-D/3-D tool paths along with S, clad overlap, an important process parameter. An STL model of an impeller part has been imported as the case study of visual simulation for the study. Table 2 is the list of process parameters and geometric parameters acquired from the case study of visual simulation for the impeller model. Fig. 7 shows a photo image taking the impeller part built by DMT process (Fig. 7(a)), a computer generated image of visually simulated impeller model for DMT process (Fig. 7(b)), and an enlarged view image of the visually simulated impeller model (Fig. 7(c)) as ordered. In the results, total period of process time is expected to be 5 hours and 23 minutes when the usage of the material is approximately 2.25 kg. In this paper an advanced visual simulation methodology as design support functionality for e-manufacturing framework has been discussed with a case study implemented since careful examination of the visually simulated model prior to the actual fabrication help minimize unwanted design iterations. In this approach, a 3-D geometric model, visually simulating the physical part to be made by DMT process, has been con- structed and utilized as the design support tool for emanufacturing, which turns out to provide some advantages in this study as in the following: First, as implementing a visually simulated 3-D model following the practical process parameters of DMT, designers can estimate the exact time spent for the process before the real fabrication (build time estimation)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000008_3352593.3352606-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000008_3352593.3352606-Figure1-1.png", + "caption": "Figure 1: Two actuated wheels mobile robot", + "texts": [ + " In section 3, the proposed event-triggered scheme with the derivation of the triggering condition is presented. Simulation results are discussed in section 4 and the paper is concluded in section 5. AIR \u201919, July 2\u20136, 2019, Chennai, India Sami Al Issa1,2, Manmohan Sharma1 and Indrani Kar1 The dynamic model of the nonholonomic mobile robot is briefly introduced in this section. The strict feedback form of systemmodel is also presented that is required for the backstepping control design. Furthermore, the control objective is stated. The systemmodel of the mobile robot shown in Fig. 1 is divided into two parts, kinematic and dynamic models. The following equations describes the kinematic model of the nonholonomic robot. \u00dbx = v cos\u03b8 , \u00dby = v sin\u03b8 , (1) \u00db\u03b8 = w . The above equations can be written in form of \u00dbq = S(q) \u03bd , (2) where S(q) = \u00a9\u00ab cos\u03b8 0 sin\u03b8 0 0 1 \u00aa\u00ae\u00ac , \u03bd = (v,w)T are the linear and angular velocities and q = (x,y, \u03b8 )T denotes the position and orientation of the robot. According to Euler-Lagrange, the following equation describes the dynamics of general class of nonholonomic mechanical systems M(q) \u00dcq +C(q, \u00dbq) \u00dbq +G(q) = E(q) \u03c4 \u2212AT (q) \u03bb , (3) where q \u2208 Rn represents generalized coordinates and \u03c4 \u2208 Rr is the input vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003282_s10015-020-00616-4-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003282_s10015-020-00616-4-Figure1-1.png", + "caption": "Fig. 1 3D CAD and the developed towed ROV not including the multi-camera unit", + "texts": [ + " Our future goal is to use the towed ROV for investigation of spatial distribution of coral reef coverage and planning of coral reef aquaculture, and investigation of topographical change caused by erosion and sedimentation. Additionally, it is useful for the management and preservation of the underwater archaeological site. For such global and/or periodic tasks, we prioritize creating wider maps using the towed vehicle with multiple cameras over creating precise maps by tradeoff in this paper. We designed and developed the human-portable towed ROV as presented in Fig.\u00a01. Its specifications are presented in Table\u00a01. Its respective length and width are 1.45 and 0.998\u00a0m, respectively. The vehicle weight with the multi-camera unit is approximately 37.5\u00a0kg. The main wings and tail wing angles 1 3 can be changed to adjust the vehicle depth and pitch angle. Horizontal tail wings play an important role in changing the vehicle pitch angle and in adjusting the angle of attack of the main wings. The NACA0015 symmetrical wing was adopted for both main and tail wings. Analytical software (XFL5 Ver" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000090_icusai47366.2019.9124863-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000090_icusai47366.2019.9124863-Figure3-1.png", + "caption": "Fig. 3. Distributed communication configuration of multi-UAV", + "texts": [ + " In the multi-aircraft collision avoidance problem, the cooperative approach is inter-aircraft communication, which aims to optimize the overall decision-making of multi-aircraft non-collision flight. The multi-UAV distributed cooperative collision avoidance architecture is shown in Fig. 2. (1) Communication between UAVs After obtaining the information of obstacles and targets, UAVs can share with UAVs within their communication distance to realize cooperative perception of the environment[16]. The communication schematic diagram between multiple UAVs is shown in Fig. 3. (2) Collaboration mechanism of multi-UAV collision avoidance system The biggest difference between multi-UAV cooperative collision avoidance system and single-UAV autonomous collision avoidance system is that UAV can realize a more comprehensive perception of the environment through interaircraft communication to guide UAV cooperative collision avoidance guidance decision-making and achieve overall optimization of decision-making behavior. 978-1-7281-5859-4/19/$31.00 \u00a92019 IEEE 2 Authorized licensed use limited to: University of Canberra" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003295_s00170-020-05693-0-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003295_s00170-020-05693-0-Figure3-1.png", + "caption": "Fig. 3 Illustration of the conical grinding milling", + "texts": [ + " The origin OG is the center of the left lateral face, and ZG is the tool axis from the left end to the right end. The XG axis and YG axis are on the left lateral face and perpendicular to each other. In the wheel coordinate system, the parametric representation of the wheel GW(h, \u03b8) can be expressed in Eq. (2): WG \u00bc h; \u03b8\u00f0 \u00de \u00bc R cos \u03b8 R sin \u03b8 h 2 4 3 5; \u00f02\u00de where h \u2208 [0, H], \u03b8 \u2208 [0, 2\u03c0]. The wheel surface normal is derived as follows: NG \u00bc h; \u03b8\u00f0 \u00de \u00bc cos \u03b8 sin \u03b8 0 2 4 3 5: \u00f03\u00de To describe the conical flank grinding processes, a tool coordinate system noted as OT is established in Fig. 3, of which the origin OT is located at the center of the machining end of the tool. As described in Fig. 3, ZT axis is the tool axis pointing from the left to the end plane, YT axis is in the vertical direction, and XT axis is horizontal. The flank grinding processes can be resolved into two steps: (1) wheel set-up and (2) wheel moving with a 4-axis archimedes spiral cord. Firstly, in the set-up operations, the grinding wheel should be configured precisely at a specific position and orientation which are shown in Fig. 3. The wheel position is defined by the wheel center OG, which is denoted by the coordinate value of [0, y0, z0]. The wheel orientation is defined as the angle between the wheel axis ZG and the tool axis ZT, which is denoted as \u03b20. The corresponding matrix of the set-up operations in the tool coordinate system can be expressed in Eq. (4): M1 \u00bc 1 0 0 0 0 cos \u03b20\u00f0 \u00de \u2212sin \u03b20\u00f0 \u00de y0 0 sin \u03b20\u00f0 \u00de cos \u03b20\u00f0 \u00de z0 0 0 0 1 2 664 3 775: \u00f04\u00de In the 4-axis grinding process, the kinematic of the grinding wheel can be resolved into 4 motions in the wheel coordinate system: (1) translation \u0394yt along YT, (2) translation \u0394zt along ZT with translating velocity v, (3) rotation about ZT with rotating speed \u03c9, and (4) rotation \u0394\u03b2t about XT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003829_s10846-020-01290-1-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003829_s10846-020-01290-1-Figure4-1.png", + "caption": "Fig. 4 The sketch of the refinement module. The black, yellow and red arrows indicate the police module output, linear movement and elastic movement directions, respectively", + "texts": [ + " Although the policy module learns the collision avoidance awareness from many human demonstrations, the learning objective is only to maximize the similarity and the collision-free property may not be fully guaranteed. Therefore, an elastic movement model is introduced according to the relative distance between needle tip and obstacle. For each sampling point pt i around the needle tip, a movement vector is calculated by, hi = \u23a7 \u23a8 \u23a9 rt Is(pt i )(pc \u2212 pt i ) ||pc \u2212 pt i ||2 , ||pc \u2212 pt i || \u2264 rt 0, ||pc \u2212 pt i || > rt (9) where rt is a radius parameter, as is shown by the blue circle in Fig. 4. rt is smaller than the receptive field of neighborhood-sampling module. It means that the elastic movement only works when the distance between obstacle and needle is small. The elastic movement ve is the sum of all the movement vectors determined by all the sampling points, namely, ve = \u2211N i=1 hi || \u2211N i=1 hi || (10) The needle tip movement is further adjusted by the weighted sum of ve and v\u2032 o, v\u2032\u2032 o = (1 \u2212 \u03c4)v\u2032 o + \u03c4ve ||(1 \u2212 \u03c4)v\u2032 o + \u03c4ve|| (11) \u03c4 = ||pc \u2212 pg||\u221a w2 + h2 (12) Finally, the relative movement of the needle tip is given by dm = dv\u2032\u2032 o (13) where d is the step length; dm is the relative movement action" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003213_j.isatra.2020.06.003-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003213_j.isatra.2020.06.003-Figure6-1.png", + "caption": "Fig. 6. Single-link arm.", + "texts": [ + " This figures also depicts the states trajectories of the investigated system under the designed controllers. Fig. 5 depicts the output under the above switching signal. Please cite this article as: S. Du, X. Li, S. Sun et al., Stability analysis and stabilization of discrete-time switched Takagi\u2013Sugeno fuzzy systems. ISA Transactions (2020), https://doi.org/10.1016/j.isatra.2020.06.003. Example 3. Now we use a single-link robot arm, borrowed from [49], to verify the efficiency of the proposed stabilization strategy. Such a single-link robot arm is illustrated in Fig. 6 with the dynamic \u03b6\u0308 = \u2212 gLOr Jr sin(\u03b6 ) \u2212 \u03c7 Jr \u03b6\u0307 + 1 Jr u + Dw where r = {1, 2} denotes the switching mode. The angle and the angular velocity of the robot arm are denoted by \u03b6 (\u03f1) and \u03b6\u0307 (\u03f1), respectively. The mass of the load is Or . Jr is the moment of the inertia, L is the value of the length, g is the acceleration of gravity, and \u03c7 is the damping coefficient. In this paper, L = 0.5, g = 9.81, \u03c7 = 2, and the mass of the load Or and the inertia Jr are presented in Table 3. Denote x1 = \u03b6 , x2 = \u03b6\u0307 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000481_ls.1256-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000481_ls.1256-Figure5-1.png", + "caption": "Figure 5. Visualisation position.", + "texts": [ + " This particle image shown in Figure 4 was only a part of the original recorded particle image, and its dimension was about 100 \u03bcm. The other part of the original recorded particle image without display was chaotic. During data-processing, an operation named as \u2018ROI Extract\u2019 (Region of Interesting) was conducted, so the dimension of the region of velocity field in the following contents was much smaller and was only several micrometres approximately. After \u2018Average Filter\u2019 processing, the velocity field looks uniformity. The visualisation position was illustrated in Figure 5, l was about 20.38 mm, w was about 36.08 mm and h was about 6.5 mm in this study. The effect of rotational speed on lubricant\u2019s velocity field during lubricating process was studied first. When the inlet pressure was 0.5 MPa, the velocity fields of different rotational speeds were Copyright \u00a9 2014 John Wiley & Sons, Ltd. Lubrication Science (2014) DOI: 10.1002/ls illustrated in Figure 6. The velocity angle was defined as the angle between the velocity vector and the vertical direction (outlet direction)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure2-1.png", + "caption": "Fig. 2. Maneuvering of the Snakeboard.", + "texts": [ + " Section 6 concludes the paper, and Appendix A summarizes the geometric and dynamic parameters of the multibody model. The Snakeboard shares the idea of forward propelling without kicking the ground. This system consists of a longitudinal bar that connects two sets of wheels, being allowed to rotate independently. In order to achieve a forward motion, the rider carries out alternate rotations with their ankles (exerting torques r and f on the rear and front wheelsets, respectively) and torso (torque \u03c8 ). These rotations are illustrated with the angles \u03c6r , \u03c6f and \u03c8 in Fig. 2 . In this way, the sequence of motion is shown in Fig. 3 . Firstly, both platforms are turned in, and the twist of the rider\u2019s torso through an angle \u03c8 allows the motion depicted by configuration (1) in Fig. 3 . Subsequently, the rider turns both feet out and moves the torso in the opposite direction, describing the motion denoted by (2). Repetition of this sequence over time leads to the forward propelling of the Snakeboard. In the case of the Waveboard, the forward propelling results from the oscillatory motion of both decks through lateral actions exerted by the rider with their feet", + " (56) for the limit case a \u2192 0 has been done, turning the torus into a ring. The obtained results coincide with those of the previous section, validating the implementation of the toroidal model. It is interesting to see the influence of the aspect ratio a b on the external actions \u03c4 . Figs. 29 and 30 show that, as the aspect ratio a b increases (moving away from the ring case a \u2192 0), the amplitude of F 2 y and F 5 y rises. These forces oscillate around zero, as occurred in the ring model (see Fig. 15 ) and as predicted in Fig. 2 of Section 2 , when the maneuvering of the Waveboard was explained. Note that this increase of amplitude is particularly important in the rear force F 2 y , becoming more similar to the forward F 5 y as a b rises. In the case of the force F 2 x , it can be seen in Fig. 31 that the rise of the aspect ratio leads to larger variations of this force. Nevertheless, this variation (around 28 N for a b = 0 . 2 ) remains small compared to the result obtained for the trajectory with constant angular velocity of Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001175_tmag.2014.2356593-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001175_tmag.2014.2356593-Figure1-1.png", + "caption": "Fig. 1. Main parameters of MD500 helicopter. (a) Top view of MD500. (b) Literal view of MD500.", + "texts": [ + "2356593 bonding to the transmission lines, and according to Maxwell\u2019s equations, the potential and electric field can be obtained by solving the governing equations shown as follows [3]: \u23a7 \u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 \u2212\u2207 \u00b7 (\u03b5\u2207\u03c6) = \u03c1 in \u03c6 = u on D \u2202\u03c6 \u2202n = q on N \u222e c \u03b5 \u2202\u03c6 \u2202n d = Qc , \u03c6 = \u03c6c on c (1) where \u03b5 is the relative permittivity, is the solve domain of potential \u03c6, \u03c1 is the charge density inside the solve domain, D and N correspond to Dirichlet and Neumann boundary, u and q are the boundary value, respectively, n is the outer normal vector, c is the boundary of floating conductor, Qc is the charge quantity of floating conductor, and \u03c6c is the potential of floating conductor. The finite element model of helicopter live-line work on 1000 kV UHV transmission lines consists of MD500, work platform, line worker, and 1000 kV transmission lines [4], [5]. Accurate models of MD500 and line worker are built in this paper. The main structure parameters of MD500 are shown in Fig. 1. The main parameters of line worker are displayed in Table I, which are mainly based on the statistical data recorded in GB10000-88 human dimensions of Chinese adults. At last, the modified data of line worker is used in consideration of the conductive clothing line worker wears. In consideration of the structure complexity of helicopter MD500 and the computer memory size, some simplifications of the helicopter are adopted, and the interconnecting pieces of different parts of the helicopter are ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001482_1.4029187-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001482_1.4029187-Figure7-1.png", + "caption": "Fig. 7 (a) A Stephenson type III six-bar linkage, (b) a dead center position of Stephenson type III six-bar linkage, and (c) a dead center position Stephenson type III six-bar linkage", + "texts": [ + " Thus, the corresponding equivalent linkage I12I23I37I71 comprised the instant centers of two neighboring links. If the three passive joint I23, I37, and I71 of the equivalent linkage lie on a common straight lie, the equivalent linkage is at the dead center positions and hence the whole linkage must be at dead center positions (Fig. 6), which can be verified by Eqs. (1) and (2). Example 6. Dead center positions for the Stephenson type III six-bar linkage. A Stephenson type III six-bar linkage contains a four-bar loop and a five-bar loop, as shown in Fig. 7. With the input given through the link 2 or the joint I23, the Stephenson type III six-bar contains two types of equivalent four-bar linkages. The type I four-bar linkage consists of the four-bar loop, or the four instant centers I12, I23, I34, and I41. The type II four-bar linkage consists of the original four links as link 1, link 2, link 5, and link 6. Thus, the corresponding equivalent linkage I12I25I56I61 comprised the instant centers of two neighboring links. When the three passive joint I23, I34, and I41 of the type I equivalent linkage lie on a common straight lie, the equivalent linkage is at the dead center 044501-4 / Vol. 7, NOVEMBER 2015 Transactions of the ASME Downloaded From: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use positions and hence the whole linkage must be at dead center positions and the three passive joint I25, I56, and I61 of the type II equivalent linkage I12I25I56I61 are collinear automatically, as shown in Fig. 7(b). If the three passive joint I25, I56, and I61 of the type II equivalent linkage I12I25I56I61 are collinear, the equivalent linkage is at the dead center positions and hence the whole linkage must be at dead center positions while the three passive joint I23, I34, and I41 of the type I equivalent linkage are not necessarily collinear, as shown in Fig. 7(c), where the collinearity of the three passive joint I25, I56, and I61 of the type II equivalent linkage I12I25I56I61 are not caused by type I equivalent linkage. Therefore, for all the dead center positions of the Stephenson type III linkage, the three passive joints of all type I equivalent four-bar linkages must be collinear or the three passive joints of such type II equivalent four-bar linkages must be collinear. These discussions also hint the special case that is these two types of equivalent linkages are at dead center positions at the same time, as discussed in Ref" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure2-1.png", + "caption": "Figure 2. Isometric view of a crawler shoe [2].", + "texts": [ + " Figure 1 illustrates the geometry of the crawler track assemblies and the crawler-formation interactions. The track is modelled using the crawler track dimensions for the P&H 4100C Boss shovel in Table 1. Only the open track chain, of the crawler assembly, in contact with the ground (Figure 1) is used for this study. Since the crawler track is made up of crawler shoes, a crawler shoe model is developed first and then connected together to form the multi-body model of the track assembly. The crawler shoe model for electric shovel 4100C Boss crawler is shown in Figure 2. This model is based on the actual crawler shoe model for the P&H 4100C Boss shovel [1,2]. The link pins are used to connect the two crawler shoes to form a continuous track. The link pin is not included in the multi-body model and will be considered in the analysis using joint constraints equations. The other crawler system components such as front and rear idlers, drive tumbler, lower rollers and guide rails are attached through joints to the crawler frame which in turn is fixed to the carbody of the shovel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002283_978-3-319-44156-6_1-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002283_978-3-319-44156-6_1-Figure1-1.png", + "caption": "Fig. 1 Loads acting on a slewing bearing", + "texts": [ + "eywords Four-contact-point bearing \u22c5 Slewing bearing \u22c5 Manufacturing errors \u22c5 Friction moment Friction moment is a very relevant parameter when designing slewing bearings. This type of bearing is used for orientation purposes in applications like tower cranes or wind turbines, where large loads are involved (Fig. 1). In most cases, the rotation is engine driven, so the friction moment must be estimated in order to dimension it. With this aim, Leblanc and Nelias [9, 10] solved the ball kinematics problem, which was later simplified for low-speed applications by Joshi et al. [7], and then used for the analytical formulation of the friction torque. This model assumes full sliding in the ball-raceway contacts, so the effects of the stick-slip regime are not considered. Recently, a Finite Element (FE) model has been I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002512_iciea.2016.7603996-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002512_iciea.2016.7603996-Figure1-1.png", + "caption": "Fig. 1 Scheme of a robotic manipulator", + "texts": [ + " For the purpose of improving contour tracking performances, it is natural to extend the known nonlinear PD control [7-13, 17] in time domain to position domain. In this paper, the master-slave synchronization of the position domain controller is combined with the flexible behaviours of the nonlinear PID controller in order to develop a new PD-type nonlinear controller in the position domain and to apply it for contour tracking of nonlinear multi-DOF robotic manipulators. Hence, this research is an expansion and further development of our previous work so as to further improve contour tracking performance [17]. A robotic manipulator, shown in Fig. 1, is used as an example to describe the concept of position domain control. In position domain control, a master-slave motion scheme is utilized. The master motion, motion 1, is used as a reference that will not introduce any error to the contour error, and the slave motions (motions 2 to 4) are described as functions of the master motion according to the contour trajectory requirements. The master motion operates in time domain, controlled by a conventional time domain controller. On the other hand, the slave motions operate in position domain, using the master motion as reference instead of time [14]. To facilitate this, the concept of a relative derivative is introduced and then the dynamic model of the robotic manipulator\u2019s slave motions are transferred from time domain to position domain, represented as functions of the master motion. Fig. 1 Scheme of position domain robotic system. 2407978-1-4673-8644-9/16/$31.00 c\u00a92016 IEEE The relative derivative concept is introduced to define the relationship of master-slave motions. For a multi-DOF robotic manipulator with n active joints, a relative derivative of slave joint s with respect to the master motion, m, is defined as [14]: s s s m m dq q q dq q \u2032 = = (1) From Eq. (1), it can be deduced that is the ratio between the velocities of the slave motion and the master motion and it describes a synchronized speed relationship between these two motions", + " (1), the second relative derivative, also called relative position acceleration, can be defined as follows: s s m dq q dq \u2032 \u2032\u2032 = (2) Expanding and rearranging Eqs (1) and (2) leads to the following: ( )2 s m s s m s m s q q q q q q q q \u2032= \u2032\u2032 \u2032= + (3) Eq. (3) show the relationship between absolute and relative motions, relating absolute velocity and acceleration in time domain to relative velocity and acceleration in position domain. Therefore, Eq. (3) constitutes an one-to-one nonlinear mapping from time domain to position domain and is used to transform the dynamics of the slave motion from time domain to position domain [15]. A dynamic model of a multi-DOF robotic manipulator, shown in Fig. 1, can be expressed as [4, 18, 19]: ( ) ( ) ( ) ( ) ( ) ( ) ( ), , ,M q q t C q q q t G q F t q q t\u03c4+ + + = (4) where ( )q t , ( )q t and ( )q t are 1n \u00d7 joint position, velocity and acceleration vectors, defined as functions of time; ( )M q is the symmetric positive-definite n n\u00d7 inertia matrix; ( ),C q q is the n n\u00d7 matrix of coriolis and centrifugal forces; ( )G q is the 1n \u00d7 vector of gravity terms; ( ), ,F t q q is the 1n \u00d7 vector of friction forces, and ( )t\u03c4 is the 1n \u00d7 vector of joint torques/forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001303_0954406213515644-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001303_0954406213515644-Figure3-1.png", + "caption": "Figure 3. Schematic diagram of sealing structure in globe valve.", + "texts": [ + "25 Ra (Ra is the profile arithmetic average error). Based on the above simulation method, a randomly rough surface with the corresponding value Ra by controlling the value of the profile height distribution function can be modelled. Figure 1 shows a modelled three-dimensional topography of metal surface with \u00bc 1 mm and x\u00bc y\u00bc 1.5, whose cross section profiles in transversal and longitudinal directions are shown in Figure 2. at MICHIGAN STATE UNIV LIBRARIES on June 26, 2015pic.sagepub.comDownloaded from Micro-contact model of metal surfaces Figure 3 shows the metal contact-sealing structure of the globe valve. From a micro-scale view, metal seal contact can be reduced to the contact between two elastic rough surfaces, and the micro-contact model is shown in Figure 4. As a consequence, the microcontact between the two flat surfaces, in fact, is regarded as the contact between asperities of rough surfaces. For a single asperity contact, the interfacial profile can be approximately taken as a spherical surface. The micro-contact state between two asperities is shown in Figure 5, where R1 and R2 are the interfacial curvature radii of two asperities, respectively, i is the mutual approach, Wi the external load and Ai the contact area for a couple of contacting asperities", + " If not, S will be modified, namely, S \u00bc S 2 when Pn i\u00bc1 Wi 4W and S \u00bc S\u00feKS0 2 when Pn i\u00bc1 Wi 5W; then repeat steps (4) and (5) if convergence is not obtained; (f) output at MICHIGAN STATE UNIV LIBRARIES on June 26, 2015pic.sagepub.comDownloaded from the approaching separation DS and calculate the maximum interfacial stress max and the ratio A/A0 of the real contact area and the nominal contact area; (g) judge whether the leaking paths are formed or not; if the leakage occurs, the external loads will increase by the step Wk\u00fe1 \u00bcWk \u00feDW and then repeat the steps from (3) to (7) until the leakage is stopped. The calculation process mentioned above is shown in Figure 11. Taking the globe valve shown in Figure 3 as an example, the micro-contact of metal surfaces is calculated and modelled. For the surface of metal contact seals machined by lapping, its root-mean-square roughness is generally in the range from 0.5 to 2 mm. In this case, three different contact surfaces are taken for example, and each roughness is 0.5, 1 and 2 mm, respectively. They all have the same sampling contact area 1 1mm and a meshed grid 200 200. Figure 12 shows the contact situations between two interfaces at the same surface roughness ( \u00bc 1 mm)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002140_icevtimece.2015.7496686-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002140_icevtimece.2015.7496686-Figure4-1.png", + "caption": "Figure 4. Prototype of PANI/Zn battery.", + "texts": [ + " After this time no more polymer could adhere to the electrode and the solution started to become green (Emeraldine Salt (ES)) indicating growth of non-adhered polymer in solution (Figure 3). FT-IR spectrum of the electropolymerized PANI was recorded with a FTIR Bruker Alpha. Thermo Scientific Evolution 220 Series UV-Visible Spectrophotometer was used to record the absorption spectrum of the electropolymerized PANI, using N,N-Dimethylacetamide (DMAC) as the solvent, with the wavelength range is 300 - 1100 nm. Then, Raman spectroscopy also used to characterize the PANI film with blue laser (488 nm). The main scheme of the PANI/Zn secondary battery was a sandwich form (Figure 4), which consisted of a filter paper sheet of porous cellulose separator, and a PANI electrode (film-type, 4 cm2) on graphite sheet as a cathode current collector. The electrolyte of batteries was an aqueous solution of 1.0 M ZnCl2 with pH 5. Triton X-100 is employed in the battery electrolyte to prevent Zn dendrite formation [12]. The PANI/Zn secondary battery was characterized by using electrochemical measurements. Electrical impedance spectra were obtained with a frequency response analyzer (Gamry Reference 3000 electroanalysis instrument), in a frequency range from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000340_jae-141878-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000340_jae-141878-Figure2-1.png", + "caption": "Fig. 2. 2D model of flux focusing magnetic gear.", + "texts": [ + " If the relationship between the steel poles is chosen to be p1 = |p3 \u2212 n2| then the rotors will interact via a common space harmonic [1,4], the angular velocities for each rotor is then related by \u03c91 = p3 p3 \u2212 n2 \u03c93 + n2 n2 \u2212 p3 \u03c92 (1) A MG does not require gear lubrication and has the potential for high conversion efficiency [3]. A MG also has inherent overload protection since if excessive torque is applied the MG will simply slip, in contrast, a mechanical gearbox would catastrophically fail. The air-gap flux density in a magnetic rotor can be substantially increased by arranging the magnets in a flux focusing arrangement. An example of a flux focusing magnetic gear (FFMG) being investigated is shown in Fig. 2 [5]. It has p1 = 4 pole-pairs, n2 = 17 steel poles and p3 = 13 pole-pairs on the \u2217Corresponding author: Kiran Uppalapati, Laboratory for Electromechanical Energy Conversion and Control, Electrical and Computer Engineering, University of North Carolina at Charlotte, 9201 University City Blvd., 28213, Charlotte, NC, USA. E-mail: Uppalapati.K.2010@ieee.org. 1383-5416/14/$27.50 c\u00a9 2014 \u2013 IOS Press and the authors. All rights reserved outer rotor. In this embodiment the magnets are magnetized along the azimuthal direction and the flux is forced into the steel poles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001104_icra.2014.6907171-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001104_icra.2014.6907171-Figure1-1.png", + "caption": "Fig. 1: Robotic Drilling System pipe handling devices", + "texts": [ + " INTRODUCTION Robotic drilling has become an increasingly desirable goal as the oil industry seeks to enhance operator safety and reduce costs. Robotic Drilling Systems AS (RDS) has developed an autonomous rig for unmanned exploration drilling in harsh environments. Contact and slip event detection is an important capability for reliable robotic manipulation, enabling quick and appropriate reaction to unanticipated events. This paper addresses unintentional slips between robotic end effectors (grippers and roughnecks) and their payloads: large, heavy metal pipes (Fig. 1). This is a challenging new robotics application for which no prior solutions exist. Numerous robotic slip detection methods exist, many of which rely on detecting minute displacements between an end effector and payload using mechanical, optical, magnetic or even thermal transducers. Bio-inspired methods often measure the vibrations or rapid changes in contact stress that accompany incipient slips. These dynamic sensing approaches have the advantage of allowing response before noticeable movement occurs [1,2]", + " If prior knowledge regarding event classes is unavailable, unsupervised classifiers can be used to perform event classification. Some methods proposed for AE data analysis include max-min distance, cluster seeking, k-means, and isodata [22,23]. In both supervised and unsupervised cases, prior work has demonstrated the ability to distinguish between different types of AE events [24]. To the authors\u2019 knowledge, such real-time AE features have not previously been used for classifying robotic contact events. The system manipulates pipes with a robotic gripper and a roughneck (Fig. 1). Both the gripper and roughneck grasp steel pipes with hardened steel dies. Contact events (including impacts, slips, and other unspecified sources of noise) may occur at either site. The robotic gripper\u2019s primary function is to transfer pipes and move them into or out of the roughneck. During pipe transfer and placement, slips may occur along the pipe\u2019s axis. However, rotational slips about the pipe\u2019s long axis are unlikely. Contact event classes for the gripper are: (I) linear slip, (V) impact, and (VI) noise (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001131_icrom.2013.6510116-FigureI-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001131_icrom.2013.6510116-FigureI-1.png", + "caption": "Figure I - Model of two wheel mobile manipulation with a reaction wheel", + "texts": [], + "surrounding_texts": [ + "The dynamics and kinematics model of a two wheel mobile manipulator robot with a reaction wheel is described in this section. For the considered robot which is shown in Figures (1-2), its parameters specifications are given in Table (1). The base of the mobile manipulator is assumed as a passive joint. Therefore, the manipulator part consists of three links and actuators in 3D workspace. The reaction wheel is placed on the first link which has one more degree of freedom rather than this one. The dynamics equations of motion of a mobile manipulator are described as: '( = M(q)q + H(q,q) + G(q) (1) where q=[ql'q2'\" .. ,qtER7 is the vector of generalized coordinates and =[\" 2,0, 4' 5' 6' lER7 is the input generalized torque and M(q) ER7x7, H(q, q) ER7xI, G( q) ER7X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively. The considered two WMM robot with a reaction wheel required complex dynamics modelling as a result of under actuated system. The passive joint causes that the balancing challenge of this robot in the XOY plane is more important than other its positions. Thus, Double Inverted Pendulum Model (DIPM) is utilized to simplify the dynamic analysis for the balancing control, [18]. The effectiveness of this simplified model is demonstrated for the dynamic locomotion of the highly nonlinear and complex system, [19]. Tablel - Parameters of the two WMM with a reaction wheel o XYZ OI_XIYIZI Q,&q2 Q3 Q;(i=4,5,6) q, L 1,(i= I ,2,3) 14 Word coordinate frame Mobile manipulator coordinate frame Rotation angles of wheels Inclination angle of the passive joint Joint angles of links Angular position of the reaction wheel Radius of wheels Distance between wheels Length of links Distance of the reaction wheel centre to first joint B. Dynamics Modelling of Double Inverted Pendulum It is possible to model two WMM as a virtual double inverted pendulum model. In this model, the components of second and third of manipulator in XOY plane are considered as the virtual second link of double inverted pendulum model. This simple model is used as the stabilizing control part. This model is shown in Figure (3) and parameters of these models is stated in Table (2). Table2 - Parameters of DIPM ilw Rotation angle of wheels ill Joint angle of the first link ile Joint angle of the second link il2 Angular position of the reaction wheel 11 Length of the first link Ie Length of the second I ink 12 Distance of reaction wheel centre to first joint xme X position of the me in 01 coordinate frame Yme Y position of the me in 01 coordinate frame Xl X component of the first link Length Y1 Y component of the first link Length me Equivalent mass on the second link m1 Mass of the first link m2 Mass of the reaction wheel mw Mass of wheels c,(i=2,e) Coefficient of friction a1 Length of the CoG on the first link i1 Moment inertia of link I around the ZI axis i2 Moment inertia of the reaction wheel around the zraxis Calculation of positions of the virtual link xmeand Yme' and its length Ie' joint angleqe and the mass meare as following: (2) (3) (4) (5) The dynamics equations of motion of this model are obtained by Euler-Lagrange equation as: (6) whereq = [qw, q1' q2' qeF ER4 is the vector of generalized coordinates and r = [fw, f1' f2' feF ER4 is the input generalized torque for the virtual double inverted pendulum modeI.M(q) ER4X4,A(q, q) ER4XI,G(q) ER4X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively which are calculated as: M = [\ufffd\ufffd\ufffd M31 M41 where: Mll = (mw + m1 + mz + me)rZM12 = M21 = (a1m1 + me (11 + Ie COS(qe)) + 11mz)rcOS(q1) + mwrZM14 = M41 = melerCOS(q1 + qe) Mzz = i1 + iz + a1 Z m1 + 11 z mz + 11 z me + Ie z me cosz qe + 211 leme cos qe + mWrZM23 = M32 = M33 = izMz4 = Ie z me cosz qe + 11 leme cos qe M13 = M31 = M34 , Z = M43 = OM44 = Ie meH1 = -fiI(mer(ft1 sin(i'iI) (\\1 + Ie COS(qe)) + leqe sin(cl!+qe)) + 11mZft1rsin(Q1) + a1 m1 ib r Sin(Q1)) - merqe(leqe Sin(q1 + qe) + leq1 COS(q1) sin(qe))Hz = -2m.leq1 (feqe COS(qe) sin(qe) + 11 qe sin(qe)) + meleqe (qwr COS(q1) COS(qe) - qe sin(qe) (\\1 + 21e cos(qe))) H3 = cZqZH4 = -me Ie rih qw COS(qe) sin(qe) + m.leq1 z sin(qe) (11 + Ie COS(qe)) + CeqeG1 = G3 = OGz = -a1 m1 Sin(q1) - 11 sin q1 (mz + me) - Ie me (cos q1 sin qe + sin q1 cos qe)G4 = -Ieme sin(q1 + qe) C. Dynamic Verification/or DIPM ADAMS is the most widely used multi-body kinematics and dynamics analysis software in the word. Also, Adams helps engineers to study the kinematics and dynamics of moving parts, how loads and forces are distributed throughout mechanical systems, and to improve and optimize the performance of their products. The extracted dynamics equations of motion of the virtual double inverted pendulum with a reaction wheel are verified by the ADAMS model which the 3D sketch of DIPM is shown in Figure (4).We insert some similar torque as input to these two models and compare the reaction of model joints with together. This torques is shown in Figure (5) which insert to the reaction wheel. Moreover, the obtained result of the model response is shown in Figure (6).These curves show that the dynamics equations of motion of the virtual double inverted pendulum are verified. So, we can use these equations in the control algorithm to realize the dynamic stability. In addition, the parameters specifications of this system in this verification routine are expressed in Table (3). 0.4 0.3 0.2 E 0.1 ;;. ! -0.1 -0.2 -0.3 I I I I I I I I I ___ 1 ____ 1___ -.l ___ .l ___ L ___ 1 __ _ 1___ -.J ___ \ufffd __ _ I I I I -- \ufffd --- 1 --- T --- \ufffd -- -0.4 - - -1- - - -1 - - - -+ - - - + - - - t- - - - 1- - - -1- - - ---j I I I I I I I I -0.5 OL--- :0c'c .5=- -L.- 1\ufffd.5=- -\ufffd2 --2cc'.C:-5-\ufffd3---=-3.L5 --':-\ufffd4\"'.5-\ufffd 5 Time (8) Since, the two WMM contains a passive joint, it requires an active controller to stabilize the passive joint and to control its stability. Related to ZMP criterion,[20], this robot have just one stable natural position that the moment of the gravity has no effect on the body, in other words, gravity force direction of COG crosses the mobile wheel axis. If we control the position of the COG to this point, we can able to satisfy the robot stability. In this section, a PID controller is proposed to control the motion of the reaction wheel to achieve dynamic stability. The supervising controller identifies the position of the COG and tunes the set point of PID controller. At the same time, the PID controller moves the reaction wheel to reach the stable natural position. Also, the control block diagram of this controller is shown in Figure (7). The supervising controller finds the COG position without any information about robot links and the related body. So, as shown in Figure (7), the input parameters of this block are the reaction wheel angular velocity and its acceleration. Therefore, it can just find the position of the COG to the right or the left from the normal position of the first link (qi=O) with this information. For an example, in Figure (2), if the reaction wheel angular velocity and its acceleration are positive, it is clear that the position of COG are in the left side from the normal position of the first link (qi=O).SO, the supervising controller changes the PID set point to the negative value. Thus, this can use many type of controller for this block such as PI, PID, Fuzzy or GA controller. In this paper we used the PI controller for this block. IV. SIMULATION RESULTS AND DISCUSSION The validation of the proposed control strategy is demonstrated by simulation results. This simulation runs at MA TLAB Simulink Toolbox for the assumed robotic system which the parameters specifications are expressed in Table(3). In addition, the controller gains are expressed in Table (4). Table4--Controller gains amount Kp Coefficient of P action 3.916 KI Coefficient of I action 1 KD Coefficient of 0 action 0.0624 Ksd Coefficient of q3 parameter -6e-3 Ksp Coefficient of q3 parameter 0.0209 This type of robot is dynamically stable and this stability is attained using a reaction wheel. Here, we consider the three bench marks and run the simulations. In the first and second one, the robot start from an initial position and make itself stable and, in third one, the robot run in the different initial position and pass the variable acceleration. The initial conditions of these case studies are respectively expressed in Table (5). Moreover, the accelerate variation curve of the third case is shown on Figure (8). Table 5 - Initial conditions for simulations Casel Casell CaselII q qp i'i2 Joint angle of the first link (deg) -30 -30 30 Joint angle of the second link (deg) -60 -120 90 Angular position of the reaction wheel (deg) 0 0 0 1.5 - - - - - I - -----1- - - - - - 1 - - --- 1 ----- - - - - - ---1 - -----1- - - - - - r- - ---- _____ --J ______ 1 ______ L ____ _ ........ 0.5 1 .\ufffd \ufffd ] -0.5 I I I ----- T -----1-----l------------ \ufffd -----r----- \ufffd----- \ufffd------------ \ufffd.5 -----\ufffd----- 4----- \ufffd------------ -20\ufffd------\ufffd,0\ufffd----\ufffd2\ufffd0 ----\ufffd3\ufffd 0------\ufffd4 0\ufffd----\ufffd 5\ufffd 0------\ufffd6 0 Time(s) Figure 8 - Acceleration variation during the operation As shown from simulation results in Figures (9-17), the time history of the robot motion is presented in Figures (9- 11) for the cases to move and reach to the stable position. Moreover, the position of the <11 must be varied until the CoG reaches to the appropriate position Figures (12-14). In this position, the robot is stable and there is no moment appears to change this style. Also, the PID modifies the set point value to satisfy the stabilization. Figures (15-17) are shown the position of the reaction wheel, during the robot reaches to stable position. These results show that the reaction wheel movement makes the internal moment that it can control the passive joint. In this control strategy, the balancing is not achieved by the robot movement and only the reaction wheel is used to stabilize. This method can also improve the dexterity of the robot motion and increases the robot controllability." + ] + }, + { + "image_filename": "designv11_34_0003406_asjc.2398-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003406_asjc.2398-Figure1-1.png", + "caption": "FIGURE 1 Mass-spring-damper system", + "texts": [ + " The nominal state-space matrix form of the servomechanism can be written as follows _X=A0X+G0U: \u00f03b\u00de The SISO mass-spring-damper system is given a demonstrated example by the proposed control method. The mass-spring-damper system is also a classical SISO servomechanism. The uncertainties of the system, including the mass of the slider, the damping coefficient, the spring coefficient, and external and friction forces are all considered in this case study. The mass-spring-damper system is illustrated in Figure 1. m0 is the known mass of the slider, \u0394m is the unknown mass of the slider, c0 is the known damping coefficient, \u0394c is the unknown damping coefficient, k0 is the known spring coefficient, \u0394k is the unknown spring coefficient, fe is the unknown external force, ff is the unknown friction force, u is the control input force, and x, _x , and \u20acx are the displacement, velocity, and acceleration of the slider, respectively. Therefore, the dynamic model can be written as M _X=NX+D+U, \u00f04a\u00de where X= x _x\u00bd T,M=M0 +\u0394M,M0 = 1 0 0 m0 , \u0394M= 0 0 0 \u0394m , N = N0+\u0394N, N0 = 0 1 \u2212k0 \u2212c0 , \u0394N= 0 0 \u2212\u0394k \u2212\u0394c ,D= 0 f e\u2212 f f\u00f0 \u00de\u00bd T ,U= 0 u\u00bd T, f f = m0 +\u0394m\u00f0 \u00deg\u03bcsign _x\u00f0 \u00de, \u03bc is the friction coefficient and sign _x\u00f0 \u00de= 1 if _x>0 0 if _x=0 \u22121 if _x< 0 8><>: : In 4a, one can find the state space matrix form _X=A0X+G0U+\u0394, \u00f04b\u00de where X= x1 x2\u00bd T, x1 = x, x2 = _x, A0 =M\u22121 0 N0 = 0 1 \u2212k0=m0 \u2212c0=m0 , G0 = 1 0 0 1=m0 , and \u0394=M\u22121 0 D+\u0394NX\u2212\u0394M _X = 0 f e\u2212 f f \u2212\u0394kx1\u2212\u0394cx2\u2212\u0394m _x2 =m0 h iT : Let a reference model of the mass-spring-damper system be _X =AmX +GmUc, \u00f05a\u00de where X = x _x \u00bd T 2R2\u00d7 1 is the desired state vector, x* and _x are the control objectives, Am and Gm are the R2 \u00d7 2 reference model matrices, Uc = 0 uc\u00bd T, and uc is the command input" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000761_ivs.2013.6629515-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000761_ivs.2013.6629515-Figure6-1.png", + "caption": "Fig. 6. Avoiding an obstacle. Static obstacle: the ordinate ys is analyzed, dynamic one: projection of ~vO is analyzed.", + "texts": [ + " The robot follows the limit cycle vector fields described by the following differential equations: x\u0307s = (sign)ys + \u00b5xs(R 2 c \u2212 x2 s \u2212 y2 s) y\u0307s = \u2212(sign)xs + \u00b5ys(R 2 c \u2212 x2 s \u2212 y2 s) (4) where (xs, ys) corresponds to the relative position of the robot according to the center of the convergence circle (characterized by an Rc radius). The function sign allows to define the direction of the trajectories described by these equations. Hence, two cases are possible: \u2022 sign = 1, the motion is clockwise. \u2022 sign = \u22121, the motion is counterclockwise. Figure 4 shows the limit cycles with a radius Rc = 1. The Obstacle is then covered by a circle, which is itself surrounded by an other virtual circle of influence with Rc radius (cf. Figure 6). The latter is chosen as the sum of the obstacle radius, the robot radius and a safety margin. \u00b5 is a positive constant. Figure 5 illustrates its influence on the limit-cycle trajectory. The choice of this constant will be rigorously discussed in section III to generate an attainable set-point. The set-point angle \u03b8Soa of the Obstacle Avoidance con- troller is given by the following relation \u03b8Soa = arctan( y\u0307s x\u0307s ) (5) The corresponding set-points (PSi , \u03b8Si ), -when the Obstacle Avoidance controller is chosen by Hierarchical Set-Point Selection block (cf", + " According to the nature of the obstacle, three cases are considered: static obstacles, dynamic obstacles, and robots of the same system. These strategies are briefly reminded in the next paragraphs. More details are available in [13]. 1) static obstacles, 2) dynamic obstacles, 3) robots of the same system. 1) Static obstacles: The same strategy proposed in [12] is maintained. Summarily, the value of sign is specified by the ordinate of the robot ys in the relative obstacle\u2019s frame (OoXoYo) (cf. Figure 6). The Xo axis of this orthonormal frame is defined thanks to two points: the center of the obstacle (which makes the origin of the frame) and the target to reach. sign = { 1 if ys \u2265 0 (clockwise avoidance) \u22121 if ys < 0 (counterclockwise avoidance) (6) The chosen direction by this strategy allows then to join the target by the side offering the smallest covered distance. 2) Dynamic obstacles: Rather than analyzing the sign of ys, it is proposed that the robot uses the obstacle\u2019s vector velocity ~vO " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003390_1350650120945517-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003390_1350650120945517-Figure2-1.png", + "caption": "Figure 2. Active incompressible lubricant film in the clearance space of two-lobe hydrodynamic journal bearing.", + "texts": [ + " \u00f03\u00de The thickness of the lubricant film in the noncircular bearing space can be determined from the following equation when the rotor center is in the static equilibrium point24 hn \u00bc 1 XJ0cos YJ0sin \u00fe 1 1 cos\u00f0 n0\u00de \u00f04\u00de where XJ0 \u00bc \"0 sin 0 and YJ0 \u00bc \"0cos 0\u00f0 \u00de are the coordinates of the static equilibrium position of the rotor center and n0 shows the angle of lobe line of centers. The parameter \u00bc Cm= C is called as preload factor. It corresponds to the amount of bearings\u2019 noncircularity and varies between 0 and 1. If preload is equal to 1.0, the lobes and bearing geometric centers coincide with each other, and noncircular bearing will be converted to the same circular type. According to Figure 2, the pressure boundary conditions are pn \u00bc 0 at \u00bc n1 \u00f05a\u00de pn \u00bc d pn=d \u00bc 0 at \u00bc n2 \u00bc n cav \u00f05b\u00de pn \u00bc 0 at z \u00bc L=2 \u00f05c\u00de where equation (5b) displays the Reynolds boundary condition applied for the solution of the governing equation, and ncav represents the position where the lubricant film of the nth lobe of a noncircular bearing becomes cavities. The static equilibrium point of rotor center and the corresponding steady-state pressure distributed in oil film can be determined by neglecting time-dependent terms of equation (2) as follows @ @ 0 h0, l\u00f0 \u00de @p0 @ \u00fe D L 2 @ @z 0 h0, l\u00f0 \u00de @p0 @z \u00bc 6 @h0 @ 0 h0, l\u00f0 \u00de \u00bc h30 12h20l\u00fe 24h30tanh h0 2l \u00f06\u00de By solving equation (6), the steady-state pressure distribution of the couple stress lubricant film can be obtained", + " To facilitate the physical sense of the values of the static and dynamic performance parameters of the analyzed two-lobe noncircular journal bearings based on the input characteristics of the journal, fluid lubricant, and the geometry of the bearings, the following table of coefficients and units is provided. In Table 2, the dimensional parameters are represented by barred symbol, and a symbol without a bar refers only to the nondimensional quantity. In this study, the noncircular two-lobe hydrodynamic journal-bearing performance is investigated for the vertical load support vector parallel to the Y axis (Figure 2). For a given value of rotor eccentricity (e), the steady-state equilibrium point of the rotor center is obtained by suitable iteration procedures.9 Furthermore, to determine the cavitation zone on each lobe, the Reynolds boundary condition equation (5b) was used for establishing the trailing edge boundaries of the positive pressure film. This condition follows from the continuity of flow at n2 demanding that the flow at n2 should be totally velocity-induced.8 Prior to investigating the effects of the couple stress parameter as well as the preload factor or noncircularity index on the steady state and dynamic performance of two-lobe journal bearings, to verify the accuracy of the developed model, the results for the pressure distribution of the circular bearing with the couple stress lubricant and the dynamic performance of a two-lobe bearing with the Newtonian lubricant are compared with the existing results from Lin26 and Lund and Thomsen27 in Figure 8 and Table 3, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003586_s12206-020-0918-5-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003586_s12206-020-0918-5-Figure1-1.png", + "caption": "Fig. 1. The 3-PPRU translational PM: (a) CAD model; (b) kinematic scheme.", + "texts": [ + " To obtain constant output velocity and force, the input velocity and force of the PM with three different Jacobian matrices are calculated. Whether they can keep constant has influence on control accuracy. In addition, a new evaluation performance index, global maximum and minimum payload mean square error index (GPSEI) to assist in evaluating payload performance, is proposed. Conclusions are drawn in the last section. The three-dimensional model and coordinate frames of the PM are established as shown in Fig. 1. It is assumed that the prismatic joints of the limb from the fixed platform to the moving platform are named as the first and the second prismatic joint, respectively. The axes of the first prismatic joint and the revolute joint are parallel, and the first prismatic joints can move along the slide rails on the fixed platform. The angles between the slide rails and their projections on the plane of the fixed platform are all represented by \u03b1 . The three projections form an equilateral triangle whose circumradius is R" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003056_5.0002805-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003056_5.0002805-Figure13-1.png", + "caption": "FIGURE 13. QUAV trajectory in a roaming flight simulation", + "texts": [ + " forward, sideward, and upward (by altitude reference). Yawing input was not demonstrated here since the yaw-control system was yet to be designed for the model. By the logic explained in Sec. 3.2, the three input-commands was translated into the fraction of thrust each rotor should give, shown in the throttle time plots in Fig. 12, along the resulted responses. Notice that the thrust can only be between 0\u20131, if beyond this limit, the input would be clipped. The QUAV simulated flight trajectory is shown in three-dimensional space in Fig. 13. Here, a noticeable characteristic of a QUAV is shown, where every time the vehicle stopped after a horizontal movement, an upward motion is resulted. This is due to the extra thrusts required for the movement before, that remain in effect when the pitch/roll angle is neutralized to stop the vehicle. The automatic altitude hold system shows an effort to reduce the upward motion by dropping the thrust of all four rotor, as shown in Fig. 12. Due to the unoptimized design, however, the thrust drop command goes beyond the rotor capability and hence resulted in a poor performance of altitude maintaining (more than 1-meter overshoot), as shown also in Fig. 13. Another noticeable dynamic of the QUAV model is the large overshoot in the roll rate (p) and pitch rate (q). One reason of this is the lack of vehicle motion effect, to the total thrust the rotors gives. For instance, the rotor physically should have a decrease of thrust when it moves upward, and an increase of thrust when it moves downward. This creates a counteracting moment whenever the vehicle rolls or pitches, and hence induce some kind of a damper. This will be a part of the model improvement in the continuation of this research" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.24-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.24-1.png", + "caption": "Figure 9.24 A Scara manipulator in its various views.", + "texts": [ + " According to this restriction, the end-effector cannot rotate about u\u20d7(4) 1 with respect to Link 1. If the specified motion of the end-effector obeys this condition, then ?\u0307?234 and ?\u0307?6 become finite but indefinite. Yet, their indefinite values become related by Eq. (9.271). Thus, this singularity induces a motion freedom in the joint space so that one of ?\u0307?234 and ?\u0307?6, say ?\u0307?6, can be assigned an arbitrary value and the other one, i.e. ?\u0307?234, can be found accordingly. Kinematic Analyses of Typical Serial Manipulators 273 A Scara manipulator is shown in Figure 9.24. Its generic name comes from the acronym SCARA that stands for \u201cSelective Compliance Assembly Robot Arm.\u201d It was developed originally under the guidance of Professor Hiroshi Makino at the University of Yamanashi in Japan. It was purposefully developed to be somewhat compliant along the first (x) and second (y) axes but definitely stiff along the third (z) axis of the base frame. Thus, it was intended to be used particularly for certain assembly operations such as inserting a pin into a hole without causing excessive lateral contact forces and moments. Normally, it is a 4-DoF (degree of freedom) manipulator without the fifth and sixth joints and its end-effector is attached to the fourth link so that u\u20d7(4) 3 is the approach vector. However, for the sake of generality, it is treated here as a 6-DoF manipulator as shown in Figure 9.24. With this general configuration, its wrist is arranged to be spherical and it is designated symbolically as 2R-P-3R or R2PR3. Its joint axes and the special kinematic details (the joint variables and the constant geometric parameters) are also shown in the figure together with the relevant unit vectors. The significant points of the manipulator are named as follows: O: Center point or shoulder point (origin of the base frame) E: Elbow point, Q: Sliding reference point, R: Wrist point, P: Tip point (a) Rotation Angles \ud835\udf031, \ud835\udf032, \ud835\udf033 = \ud835\udeff3 = 0, \ud835\udf034, \ud835\udf035, \ud835\udf036 Five of the rotation angles are joint variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002879_s10878-020-00533-z-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002879_s10878-020-00533-z-Figure6-1.png", + "caption": "Fig. 6 SCS trajectories are partitioned into rings. a two rings; b one ring; c three rings", + "texts": [ + " The second subsection presents two new concepts, tokens and token graph. This subsection presents (with no proof) some results and concepts related to rings that were introduced in Bereg et al. (2018). We enunciate them here in order to make this paper self-contained. Definition 6 (Ring) A ring in an SCS with communication graph G, is the locus of points visited by a robot following the assigned movement direction in each trajectory and always shifting to the neighboring trajectory (in G) at the corresponding link positions. Figure 6 shows various SCSs with different numbers of rings, each ring shown in a different color and stroke type. Remark 3 Each point in a trajectory belongs to a single ring, so, the rings in an SCS are pairwise disjoint and partition the trajectory points into equivalence classes. Each ring is a closed path composed of segments of trajectories and has a direction of travel determined by the movement in the participating trajectories. Definition 7 (Path in a ring) A path in a ring r from a point p \u2208 r to a point q \u2208 r is the ordered set of visited points from p to q following the travel direction of r (it may contain tours on r ). As suggested by the examples in Fig. 6, for any given system, different rings may have different. In discussing the length of a ring, it is convenient to ignore the effect on distance arising from shifting between neighboring trajectories and we assume neighboring circular trajectories are tangent to each other. Definition 8 (Length of a ring) The length of a ring is defined as the sum of the lengths of trajectory arcs forming the ring. The length of every ring in an SCS is in 2\u03c0N (Corollary 13 of Bereg et al. (2018)). Remark 4 We can extend the concept of length to a path in a ring as the sum of the lengths of trajectory arcs forming the path", + " In this section we show how to compute efficiently this measure when the communication graph is a cycle or a grid. Lemma 10 Let G be the communication graph of a m-partial SCS. If G is a cycle, then the system has exactly two rings, one with CW direction and the other with CCW direction. Furthermore, every edge of G corresponds to a crossing of the two rings. Proof We proceed by induction on the number 2M of trajectories (nodes) in G (recall that G is bipartite, so every cycle has even length). If M = 2, we have 4 trajectories, and the claim holds, as shown in Fig. 6a. Assume as inductive hypothesis that, for a fixed value M , the claim holds for every cycle graph with 2M trajectories. Now, let us consider a cycle graph G with 2(M + 1) trajectories (Fig. 16a). If we remove two consecutive trajectories (C and D) and \u201cglue\u201d the ends in the broken section we obtain a cycle graphG \u2032 with 2M trajectories (Fig. 16b). By the inductive hypothesis, the claim holds for G \u2032. Figure 16b shows the clockwise ring using solid stroke in red, and the other one using dashed stroke in blue" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002983_b978-0-12-818546-9.00016-6-Figure11.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002983_b978-0-12-818546-9.00016-6-Figure11.4-1.png", + "caption": "Figure 11.4 Smart wheelchairs: (A) intelligent wheelchair [74]; (B) smart wheelchair [75]; (C) smart wheelchair [76]; (D) navigation-system-equipped smart wheelchair [77]; (E) robotic wheelchair [78,79].", + "texts": [ + " Joysticks can be used in powered wheelchairs as an interface system to transmit the user-desired orientation to the device. For people with severe hand difficulties, there are other options as an interface system such as facial moving, chair moving and verbal actions. Smart wheelchairs refer to the powered wheelchairs equipped with different types of controlling systems to enable a wide range of patients suffering from severe physical and neurological disorders to be moved with a high level of safety. An intelligent wheelchair was designed for people having mobility disorders (Fig. 11.4A) [74]. In this smart wheelchair, a 312 Chapter 11 Assistive devices for elderly mobility and rehabilitation: review and reflection set of laser sensors was utilized to detect narrow passages and to manoeuver the device through. In addition, the wheelchair is equipped with a wall-following feature. Quaglia et al. [75] developed a smart wheelchair that enables the user to move in different conditions, climb the stairs, and to pass over obstacles easily. Four-linkage mechanism was employed to change the position of the seat in order to improve Chapter 11 Assistive devices for elderly mobility and rehabilitation: review and reflection 313 the safety and stability of the wheelchair during different conditions such as stair ascending, descending, obstacle passing, and normal using. Lopes et al. [76] equipped a robotic wheelchair with assistive navigation system for people with high level of motor disabilities and muscular weakness including amyotrophic lateral sclerosis and cerebral palsy patients (Fig. 11.4C). In this regard, a brain computer interface was used to detect the user\u2019s intentions from electroencephalography signals coming from the user\u2019s brain to navigate the device to the desired orientation. On the other hand, De La Cruz et al. [77] designed a robotic wheelchair with a navigation system for people with severe muscles\u2019 disabilities for indoor usages (Fig. 11.4D). Radio frequency robotic identification and inductive sensors were employed to follow metallic landmarks and radio frequency tags on the ground in order to navigate the wheelchair. This device is also equipped with obstacle avoidance system to improve the user\u2019s comfort. A robotic wheelchair (Fig. 11.4E) has been designed to support people suffering from mobility disorder [78]. The device works under three operation modes\u2014stop, semiauto, and manual\u2014which can be detected using a webcam by the user facial movements such as eye blinking and head shaking. A laser sensor and a Microsoft Kinect sensor have been used to detect objects and obstacles. An inertial measurement sensor has been mounted on the device to transmit inertial parameters to the controller for better control of heading direction. Ghani et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002247_jsee.2016.00070-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002247_jsee.2016.00070-Figure1-1.png", + "caption": "Fig. 1 Definition of impact angle", + "texts": [ + " (vii) The relative distance R between the missile and the target can be measured and its rate R can be calculated by the system. Considering a planar engagement between the missile and the non-maneuvering target, we propose a novel guidance law with impact angle constraint. The impact angle control guidance laws are designed to achieve interception with a desired impact angle. Here, we consider the impact angle, denoted by imp\u03b8 , is the angle be- tween velocity vectors of the missile and the target as in [21]. The impact angle geometry can be seen in Fig. 1. Denoting the heading angles of the missile and the target at the terminal time as \u03b8Mf and \u03b8Tf, respectively, the impact angle, \u03b8imp, is defined as \u03b8imp= \u03b8T f \u2212\u03b8Mf. (1) Here, for the case of stationary targets, we define the heading angle of the target as \u03b8Tf= 0. Consider a planar engagement between the missile and the slowly moving target as shown in Fig. 2, where (xM, yM) and (xT, yT) denote the position of the missile and the target, respectively. The missile heading angle, velocity, and lateral acceleration are expressed by \u03b8M, VM and AM, respectively", + " To improve the terminal performances, we consider an additional terminal constraint as follows: the time-derivative of acceleration should be zero at the terminal time, ( )M fA t = 0. Under some assumptions detailed in the beginning of this section, we describe the non-linear kinematics based engagement dynamics as cos cosT T M MR V V\u03b7 \u03b7= \u2212 (2) sin sinT T M MRq V V\u03b7 \u03b7= \u2212 + (3) / , 0M M M TA V\u03b8 \u03b8= = (4) , .M M T Tq q\u03b7 \u03b8 \u03b7 \u03b8= \u2212 = \u2212 (5) Here, the guidance law tries to force the LOS angle rate q to zero. To derive the relation between impact angle imp\u03b8 and LOS angle q, we assume that the missile and the target lie on the collision course, the geometry of which is shown in Fig. 1. At the collision course, the LOS angle rate q is zero. Thus, according to (3) and (5), we can derive the corresponding LOS angle, denoted as qf, as follows: sin sinM Mf T TfV V\u03b7 \u03b7= (6) ,Mf f Mf Tf f Tfq q\u03b7 \u03b8 \u03b7 \u03b8= \u2212 = \u2212 (7) which from (3), corresponds to that when Rq = 0. Thus, we can obtain the LOS angle at the interception course, qf, which can be given by sin arctan( ) cos imp f Tf imp T M q V V \u03b8 \u03b8 \u03b8 = \u2212 \u2212 (8) where, qf and \u03b8imp are the LOS angle and impact angle at the time of interception, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003678_ecce44975.2020.9235946-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003678_ecce44975.2020.9235946-Figure3-1.png", + "caption": "Fig. 3. Operation principle of magnetic suspension.", + "texts": [ + " In the proposed motor, the high speed rotor is magnetically suspended using the two-axis actively position regulated bearingless motor principle [8]-[11]. In this magnetic 279 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on May 17,2021 at 01:16:36 UTC from IEEE Xplore. Restrictions apply. suspension, only the x- and y-axis positions of the rotor are regulated, and the other z-axis, tilting \ud835\udf03\ud835\udf03\ud835\udc5a\ud835\udc5a and \ud835\udf03\ud835\udf03\ud835\udc66\ud835\udc66 are passively stabilized by using the attractive force caused by permanent magnets. In Fig. 3, the principle of suspension force generation in the proposed motor is shown. Let us suppose the rotor is centered, with zero suspension current, the magnetic flux density of inner airgap is symmetrical, and the amplitude of magnetic flux density is identical at points 1 and 2. Therefore, the attraction forces cancel each other, so that the force on the rotor is zero. With positive current in suspension winding, the amplitude of magnetic flux density at point 1 is increased because the direction of 2-pole flux and 4-pole suspension flux is the same" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000601_j.jsv.2014.08.016-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000601_j.jsv.2014.08.016-Figure1-1.png", + "caption": "Fig. 1. Floating frame of reference formulation.", + "texts": [ + " (20) and (22), a fact, which one can easily observe in practice when performing computations by hand, e.g. for working out some tutorial problems [4]. The approximate relations stated in Section 3 above now will be applied in the framework of the Floating Frame of Reference Formulation, which is suitable for describing the motion a deformable body having additional rigid body degreesof-freedom, and as such forms an important basis of the Multibody Dynamics technique, see Shabana [10]. In the framework of the Floating Frame of Reference Formulation, see Fig. 1, we describe the deformation of the body with respect to a rigid reference configuration, which moves together with the actual motion of the deformable body, and which is called the floating reference configuration. We rigidly attach to the floating reference configuration a Cartesian coordinate system with base vectors E - \u03b1; \u03b1\u00bc 1;2;3, denoted as the floating coordinate system, since it moves together with the floating reference configuration. The actual position vector of the origin A of the floating coordinate system, taken from the origin O of the global inertial frame, is expressed as r - A \u00bc \u2211 3 \u03b1 \u00bc 1 x\u03b1A t\u00f0 \u00de e-\u03b1 (26) For the global inertial frame, a space-fixed inertial Cartesian coordinate system with base-vectors e - \u03b1; \u03b1\u00bc 1;2;3; and with origin in point O is used in Eq", + " Since the base vectors e - \u03b1 do not change with time, we obtain the following functional representation: r - A \u00bc \u2211 3 \u03b1 \u00bc 1 q\u03b1 t\u00f0 \u00de e-\u03b1 \u00bc r - A q1 t\u00f0 \u00de; q2 t\u00f0 \u00de; q3 t\u00f0 \u00de (28) In the floating coordinate system, the directed vector from the origin A to the place of the particle in the reference configuration can be written as p -\u00bc \u2211 3 \u03b1 \u00bc 1 P\u03b1 E - \u03b1 (29) where the components P\u03b1, the Cartesian coordinates of particles in the reference configuration, do not depend on time, while the base vectors E - \u03b1 do move with the floating reference configuration. The components P\u03b1 shall be used for labeling the place of the particles in the moving reference configuration. In Eq. (4) and in the following, this is performed by introducing the time-independent vector, see Fig. 1: P - \u00bc \u2211 3 \u03b1 \u00bc 1 P\u03b1 e - \u03b1 ) Div\u03a0e \u00bc \u2211 3 \u03b1 \u00bc 1 \u2202\u03a0 \u2202P\u03b1 e - \u03b1; Grad \u2202 r\u2202qk \u00bc \u2211 3 \u03b1 \u00bc 1 \u22022 r\u2202P\u03b1\u2202qk\u2297 e - \u03b1; etc: (30) see Eqs. (1) and (23). Now, each of the moving base vectors E - \u03b1 in Eq. (29) can be obtained from the global base vectors e - \u03b1 via a single timedependent rotation tensor Re, which describes the rotation of the co-moving rigid reference configuration with respect to the global coordinate system. As is well known, the rotation tensor is orthogonal, Re 1 \u00bc ReT , and it can be completely described by three time-dependent rotation angles \u03b81, \u03b82 and \u03b83, Shabana [10]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000856_tmag.2014.2362515-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000856_tmag.2014.2362515-Figure4-1.png", + "caption": "Fig. 4. Applied stress distribution in the stator core. (a) Non-stress. (b) Model 001 (uniaxial stress). (c) Model 002 (biaxial stress). (d) Model 003 (uniaxial and biaxial stress).", + "texts": [ + "2 T, and exciting current is 5 A, and the speed of revolution is 1500 r/min. A non-oriented electrical steel sheet is used as the magnetic material of the stator and rotor cores. We try to reduce the magnetic power loss of the motor core by applying the tensile stress. The uniaxial tensile stress is applied to the region, which is magnetized under the alternating magnetic flux condition. The biaxial tensile stress is applied to the region, which is magnetized under the rotating magnetic flux condition. Fig. 4 shows the assumed stress distribution in part of the stator core. The uniaxial tensile stress at \u03c3x = 20 MPa and \u03c3y = 0 MPa is applied along the longitudinal direction on the teeth and along the circumferential direction on the core back, as shown in Fig. 4(b) and (d). The biaxial tensile stress at \u03c3x = 20 MPa and \u03c3y = 20 MPa is also applied to the root and edge of the teeth in, as shown in Fig. 4(c) and (d). The uniaxial and biaxial stress is applied to the stator core, as shown in Fig. 4(d). In addition, the SCES vector magnetic characteristic analysis and the measurement database of the vector magnetic properties under stress are used in this numerical simulation. Fig. 5 shows the calculated local vector magnetic properties of a part of the stator core in each model. The local magnetic properties at the points are evaluated, as shown in Fig. 4(a). The magnetic anisotropy in the motor core is induced by applying the stress. The local loci of Bx \u2212 By and Hx \u2212 Hy and the hysteresis loops of the x and y components with the assumed stress differed in comparison with those of the nonstress. The rotating magnetic flux in the central part of the teeth occur as shown in Fig. 5(b), and the alternating magnetic flux in the edge of the teeth and core back occur, as shown in Fig. 5(a) and (c). The change of the loci of Hx \u2212 Hy became larger than that of Bx \u2212 By " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000972_amm.711.27-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000972_amm.711.27-Figure2-1.png", + "caption": "Fig. 2 Contours of absolute pressure (Unit: Pa), temperature (Unit: K) and vapour volume fraction of (a-c) water-lubricated bearing and (d-f) oil-lubricated bearing", + "texts": [ + "48 J/(kgK) and thermal conductivity is 50 W/(mK). Since the bearing is fully submerged, the pressure at the inlet and outlet boundaries is taken as 101 325 Pa. The outer surface of the flow model is modeled as a \u201cstationary wall\u201d and the inner surface is modeled as a \u201cmoving wall\u201d with the rotational speed. A no-slip condition is imposed on the two walls. Results and discussion. In this study, flow models with different bearing lengths are calculated to find the length corresponding to the load. Results are shown in Table 2 and Figure 2. From Table 2, it can be seen that the water-lubricated bearing length is 80 mm, but the oil-lubricated bearing length is only 18.2 mm, which means that water-lubricated bearing needs to be 3.4 times longer than oil-lubricated bearing for the same load. However, the friction coefficient of water-lubricated bearing is only 0.0024, just 11.7% of that of oil-lubricated bearing. As a result, the power loss of water-lubricated bearing is much lower than that of oil-lubricated bearing. The maximum temperature rise of water-lubricated bearing is only 1.95 K due to the low viscosity and high specific heat of water, far less than 96.76 K corresponding to oil lubrication. In Figure 2, the vertical direction is the eccentricity direction. The attitude angles of water-lubricated bearing and oil-lubricated bearing are 52.5\u00b0 and 42.4\u00b0, respectively. From the pressure distribution, it is shown that as the oil-lubricated bearing is shorter than water-lubricated bearing, the pressure is much larger that of water-lubricated bearing. The maximum temperature occurs around the eccentricity direction. Cavitation occurs at the divergence region. This study quantitatively compares the lubrication performance of water-lubricated bearing and oil-lubricated bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure3.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure3.1-1.png", + "caption": "Fig. 3.1 Reference frames associated with 3D free moving rigid body", + "texts": [ + " This means the use of Galileo transformations to describe the displacement of a body with respect to a given reference frame, and a rotations from which both linear and angular velocities can be derived. These expressions are needed in the Newtonian formulation since the dynamic expression needs the inertial expression (with respect to an inertial frame) of the first and the second time derivatives of the position for every single particle in a body. The position of a frame is defined as the position of its origin w.r.t. the origin of a reference frame, expressed with coordinates of the reference frame. For the example on Fig. 3.1 is d, where g is the origin of frame 1. As for the attitude or orientation, the set of variables that describes this is not a unique expression as in the case of position.As it will be shown in the next sections, the attitude variables can be parameterized with at least three parameters, but can also be expressed with more than those. The attitude parameter vector is described in following section of the Rotation Matrix. 3.1 Translations Let 0 be a right-handed reference frame defined with the unit vectors i , j and k along the Cartesian axes x0, y0 and z0 of 0", + "4) Boundedness of the Overall Structure \u2016H(q)q\u0308 + C (q, q\u0307) q\u0307 + g(q)\u2016 \u2264 c0 + c2 \u2016q\u0307\u20162 + c3 \u2016q\u0308\u2016 (7.5) \u00a9 Springer Nature Switzerland AG 2019 E. Olgu\u00edn D\u00edaz, 3D Motion of Rigid Bodies, Studies in Systems, Decision and Control 191, https://doi.org/10.1007/978-3-030-04275-2_7 307 308 7 Lagrangian Formulation 7.1 Direct Lagrangian Expression For rigid solids, its configuration is represented by the position d = d(0) \u2208 R 3 and the orientation R(\u03b8) \u2208 SO(3) of a rigid reference frame attached to the body (see Fig. 3.1). The attitude or orientation can be parameterized with a reduced set of attitude parameters \u03b8 \u2208 R m , whose dimension m = {3, 4}, cannot be smaller than the 3 degrees of freedom of attitude, but that may be 4 for redundant attitude representations (refer to Chap. 4), where each set is known as a attitude representation. Since the pose (6.5) is defined as the vector expression of both position and attitude of a non-inertial frame x = (dT ,\u03b8T )T \u2208 SE(3), it seems natural to chose this as the generalized coordinates for a rigid body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003430_lra.2020.3015463-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003430_lra.2020.3015463-Figure1-1.png", + "caption": "Fig. 1. (a) Tracked vehicle \u201cQuince\u201d climbing up spiral stairs using reaction force from safety wall and (b) target spiral staircase with safety wall.", + "texts": [ + " The asymmetrical ground shape causes the slippage of the left or right track when the robot rotates on the spiral stairs, which complicates its rotation on the terrain. Previously, the authors experimentally demonstrated that the wall-following motion was effective for stabilizing and accelerating spiral stair climbing motion [1]. A side wall is often attached to the handrail to prevent tools or components from collapsing. A robot follows the wall with its passive wheels attached to the side surface of the sub-tracks, as shown in Fig. 1(a). Rotational motion is generated using the geometrical constraint of the wall. In our previous study, we demonstrated the improved performance of a manual operation in terms of speed using a wall-following motion on spiral stairs. In addition, we experimentally demonstrated that the method is available in several initial states with different entry angles to the wall. The wall-following motion of robots is similar to human inspectors grasping handrails during stair climbing. Through collisions with surrounding objects, motion is stabilized and certainty is guaranteed", + " David et al, proposed the turning motion on sandy soils for tracked vehicles that use plows [13]. This method enables a rotational motion on sandy ground by fixing the rotational center using a plow. This method is the same as our proposed method in that the turning motion is generated by fixing the rotational center using geometrical constraints. The sensor reflective approach is used for climbing spiral stairs because it can realize a stable motion generation even in a severe environment where slippages occur in the main-tracks or sub-tracks [6], [7]. Fig. 1(a) shows our target tracked vehicle, which has six degrees of freedom (DoF) (two main tracks and four subtracks). The six DoF motions were autonomously determined using the sensor-reflective approach. A two-dimensional (2D) LiDAR sensor was used to measure the spiral stair shape. An internal measurement unit (IMU) sensor was used to measure the pose of the main body. Four encoder sensors were used to measure the sub-track angle, and six DoF motions were autonomously determined from these sensor data in real time", + " \u03b8max = arctan ( l1 l2 ) (6) If \u03b8 > \u03b8max, the sign of denominator of \u03b8\u0307 (R(\u2212l1 cos \u03b8 + l2 sin \u03b8)) is inverted and \u03b8 diverges. However, in our target environment, it was physically impossible to achieve \u03b8max because of the narrow passage. We conducted two evaluations in this study. First, the wall following motion was effective for spiral stair climbing using the geometrical model described in Section V-A. Next, the performance of autonomous spiral stair climbing was compared with that of a manual operation. In this experiment, tracked vehicle \u201cQuince\u201d in Fig. 1(a) is used. The body parameters in Table I are as follows: l1 = 0.35 [m/s], l2 = 0.25 [m/s]. The weight of the robot is 33 kg and it has 10 kg payload. The maximum linear velocity is 0.5 [m/s]. Experiments are conducted on the spiral stairs in Fig. 1(b), which has 0.86 [m] radius and 0.23 [m] height steps. A. Verification of Trajectory Convergence During Wall-Following Motion In Section VI-A, the trajectory estimation by the model is compared with the actual robot\u2019s trajectory, and the convergence of the entry angle \u03b8 is shown from the experimental results. Robot trajectories are estimated with the initial entry angle of 30\u25e6 by measuring the robot pose using a motion capture camera on a spiral staircase. To demonstrate the convergence of the entry angle to the wall, time change of the entry angle \u03b8 is measured using the motion capture information and the wall shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002875_s42496-020-00033-7-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002875_s42496-020-00033-7-Figure5-1.png", + "caption": "Fig. 5 Particular of the sloshing equivalent mechanical model in MSC Adams", + "texts": [ + "85 h \ud835\udf11 ) mf ( 3\ud835\udf112 16 + h2 12 ) \u2212 m0h 2 0 \u2212 3 \u2211 i=1 mih 2 i , if h \ud835\udf11 < 1 1 3 The spacecraft is hereby represented as a parallelepiped platform with solar panels as appendages (respectively, in green and in red in Fig.\u00a04). The two tanks are modelled as bodies with assigned mass and inertia to reproduce a hollow cylindrical tank. The flexibility of the two solar panels is simulated using FEM techniques. In particular, the two arrays are shell bodies meshed with Quad4 elements. The two tanks are rigidly connected to the platform to simulate their fastening to the satellite. A detailed image of a single tank is provided in Fig.\u00a05. The sloshing masses ( m1 in green, m2 in blue and m3 in red) have been modelled as\u00a0spherical bodies with assigned mass deriving from Eq.\u00a0(4) and negligible inertia. Their location with respect to the centre of the filled tank respects what expressed by Eq.\u00a0(5). Such masses have been connected via translational joints to the walls of the tank, so that they are constrained to move only in the horizontal direction to simulate the lateral sloshing behaviour. At the same time, two spring\u2013damper forces are acting on each mass to reproduce the elastic forces they are subjected when moving from their equilibrium position (these forces are representative of the \u201celastic\u201d properties of the fluid)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure13-1.png", + "caption": "Fig. 13 Cutting region when peripheral cutting edge takes part in machining and yB>R1", + "texts": [ + " (15 to 16), and R1 is, R1 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi yA2 \u00fe h Rc\u00f0 \u00de2 q \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi e R ap tan \u03b1 2 h i2 \u00fe h Rc\u00f0 \u00de2 r (1) If yBR1, the cutting region can be divided into four different parts, as shown in Fig. 13. In this case, part I, part II, and part III are all the same as they are in Fig. 11. So the corresponding swept area si of the three parts can be calculated by Eqs. (31), (37), and (33). While part IV can be treated the same as part IV in Fig. 10, we can use Eq. (34). Based on the analysis above, swept area si of any IEi on the cutting edge can be calculated according to Eqs. (31) to (39). Although cutter wear cannot be avoided in turn-milling, it can be balanced on whole cutting edge if the cutting parameters are optimized, thereby prolong cutter life and reduce the cost on cutter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001131_icrom.2013.6510116-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001131_icrom.2013.6510116-Figure9-1.png", + "caption": "Figure 9 - Robot motion time history forCASE I", + "texts": [], + "surrounding_texts": [ + "The stabilization of a two Wheel Mobile Manipulator (WMM) was proposed using a reaction wheel. A virtual inverted pendulum was modelled by a 4-DOF manipulator, and it was controlled to achieve the robot balancing. The CoG position of the robot was identified by observing the reaction wheel position. By angular velocity and acceleration of the reaction wheel, the set point of PID controller was changed to the <11 position using the supervising control. This control algorithm was performed to achieve the CoG on the desired location and make the robot stable. To decrease the control calculation, this method was never used the calculation of the CoG identification. However, the supervising control method is used. Finally, simulation results proved the effectiveness of reaction wheel employing. This simulation shown that robot could start in any position and made itself stable. 4ur-------------,I---------------,I--------------\ufffdI I 30 k- - - - - - - - \ufffd - - - - - - - - - - - - \ufffd - ----------- .. ... \u00b7--\u00b7\u00b7\u00b7-\u00b7\ufffd\u00b7 ..... \u00b7r\u00b7-\u00b7-\u00b7 I 20 : - - - - - - - - - - - \ufffd - - - - - - - - - - - - \ufffd - ----------- 4U,-------,-------\ufffd------_,--------,_--\ufffd==\ufffd======\ufffd 30 ______ :______ \ufffd ______ \ufffd ______ :____ I =:- \ufffd:\ufffd\ufffd:;\ufffd\ufffdIe l 20 , , , -------------------------, , , 10 . - - - - - _1 - _____ -.J ______ l ______ 1 _______ 1 _____ _ E o - - - - - -1- - - - - - ---I - - - - - - \ufffd - - - - - - \ufffd - - - - - -1------ \ufffd -10 - - - - - -1- - - - - - ---t - - - - - - t- - - - - - - r- - - - - - -1------ -20 - - - - - -1- - - - - - I - - - - - - T - - - - - - r - - - - - -1- ----- I I I I I -30 \\- -----:- - - - - - \ufffd - - - - - - \ufffd - ----- :-------:------ '\" ____ ,I\"r ,'- ' , , -40 ---\ufffd------\ufffd \ufffd------:r--\"\"'''''''''''''''l -500\ufffd------,\ufffd0\ufffd----\ufffd2\ufffd0-------3\ufffd0\ufffd----\ufffd4\ufffd0-------=50\ufffd----\ufffd6' Time(s} Suppression of Two-Wheel Mobile Manipulator Using Resonance Ratio-Control-Based Null-Space Control\", iEEE Transactions on industrial Electronics, 57(12), pp. 4137-4146, (2010) [2] Nguyen, H. G., Morrell, 1., Mullens, K., Bunneister, A., Miles, S., Farrington, N., Thomas, K., and Gragee, D. W. \"Segway Robotic Mobility Platform\", Proc. SPiE Mobile Robots, Philadelphia, (2004) [3] Huang, Ch. and Fu, Li., \"Passivity based control of the double inverted pendulum driven by a linear induction motor\", Proc of iEEE Conference on control Applications, 2, pp. 797-802 (2003) [4] Grasser, F., D' Arrigo, A., Colombi, S., and Rufer, A. c., \"JOE: A Mobile, Inverted Pendulum\", iEEE Trans. on industrial Electronics, 49(1), pp. 107-114, (2002) [5] Niki, H. and Toshiyuki, M., \"An Approach to Self Stabilization of Bicycle Motion by Handle Controller\", Proc of the first Asia international Symposium on Mechatronics, (2004) [6] Rubi, J., Rubio, A., and Avello, A., \"Swing-up control problem for a self-erecting double inverted pendulum\", iEEE Proc. Control theory Application, 149(2), pp.169-175 (2002) [7] Liu, D., Gue, W., Yi, 1., and Zhao, D., \"Double-Pendulum-Type Overhead Crane Dynamics and Its Adaptive Sliding Mode Fussy Control\", Proc. Of the Third into Conf. on Machine Learning and Cybernetics, pp.26-29, (2004) [8] Pathak, K., Franch, J., and Agrawal, S. K., \"Velocity and position control of a wheeled inverted pendulum by partial feedback linearization\", iEEE Trans. RobotiCS, 21(3), pp. 505-513, (2005) [9] Zhang, M. and Tran, T., \"Hybrid Control of the Pendubot\", iEEE/ASME Trans. Mechatronics, 7(1), pp. 79-86 (2002) [10] Ibanez, C. A., Frias, O.G., and Castanon, M. S., \"Lyapanov-based controller for the inverted pendulum cart system\", Nonlinear Dynamics, 40(4), pp. 367-374, (2005) [11] Ge, S.S., Hang, C. c., and Zhang, T., \"A direct adaptive controller for dynamic system with a class of nonlinear parameterizations\", Automatica, 35, pp. 741-747 (1999) [12] Brooks, R., Aryanada, L., Edsinger, A., Fitzpatrick, P., Kemp, C. c., Reilly, 0., Torres-jara, E., Varshavskaya, P., and Weber, J., \"Sensing and manipulating built-for human environments\", int. J. of Humanoid RobotiCS, 1(1), pp. 1-28 (2004) [13] Miyashitaa, T. and Ishiguroa, H., \"Human-linke natural behavior generation based on involuntary motions for humanoid robots\", Robotics Autonomous Systems, 48, pp. 203-212 (2004) [14] Gans, N. R., and Hutchinson, S. A., \"Visual servo velocity and pose control of a wheeled inverted pendulum through partial-feedback linearization\", Proc. of iEEE/RSJ international Conference on intelligent Robots and Systems, pp. 3823-3828 (2006) [15] Yamamoto, Y. and Yun, x., \"Coordinating Locomotion and Manipulation of Mobile Manipulator\", Proc. of the 31st iEEE Conference on Decision and control, pp. 2643-2648 (1992) [16] Yamamoto, Y. and Yun, x., \"Control of Mobile Manipulators Following a Moving Surface\", Proc. of iEEE international Conference on robotics and Automation, 3, pp. 1-6 (1993) [17] Seraji, H., \"An On-line Approach to Coordinated Mobility and Manipulation\", Proc. of iEEE international Conference on Robotics and Automation, 1, pp. 28-35 (1993) [18] Acar, C. and Murakami, T., \"Multi-Task Control for Dynamically Balanced Two-Wheeled Mobile Manipulator through Task-Priority\", iEEE international Symposium on industrial ElectroniCS, pp. 2195- 2200 (2011) [19] Sasak, K. and Murakami, T., \"Pushing operation by two-wheel inverted mobile manipulator\", iEEE international Workshop on Advance Motion Control, pp. 33-37, (2008) [20] Alipour, Kh. And Moosavian, S. Ali. A., \"How to Ensure Stable Motion of Suspended Wheeled Mobile Robots\", Journal of industrial Robot, 38(2), pp. 139-152 (20 I I) 50,--------------.--------------,--------------. , , o ----------- t- ----------- t- ----------- -50 - - - - - - - - - - - L ___________ L - ---------- \ufffd -100 :3- , , - - - - - - - - - - -1 -- - - - - - - - - - T - ---------- -150 <>: - - - - - - - - - -1- - - - - - - - - - - - +- - ---------- \ufffd '0 -200 -250 r ' , , - - - - - - - - - -1 -- - - - - - - - - -1 - ---------- -300L--------------L--------------\ufffd------------\ufffd o 10 15 lime (s) Figure 15 - Position of the reaction wheel 100 .--------------.--------------.--------------, o -----------\ufffd-----------\ufffd----------- ___________ L ___________ L __________ _ -; -200 - - - - - - - - - - - f- - - - - - - - - - - - f- ----------- c;, c: ..: -300 -400 , , -------1 -----------1 ----------- - - - - - - - ----- -\ufffd- - --\ufffd - - - - - -----\"\"\"'- -500 '--____________ L-____________ -'-____________ --' o 5 10 15 lime (5) -50 0 \ufffd- - - ,\ufffd0- - - \ufffd20\ufffd- - \ufffd 3\ufffd0- - - \ufffd 40\ufffd- - \ufffd5\ufffd0 - - - \ufffd60 Time(s) Figure 17 - Position of the reaction wheel" + ] + }, + { + "image_filename": "designv11_34_0003363_j.mechmachtheory.2020.104027-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003363_j.mechmachtheory.2020.104027-Figure3-1.png", + "caption": "Fig. 3. Simple model of chain entering pulley.", + "texts": [ + " The bending characteristics of the chain used in this study change according to the crosssectional curve \u03b6 p 1 of pin A. This is because the cross-sectional curve of pin B is a straight line parallel to the Y k axis. A later section will explain how this characteristic affects the vibration of the chord. Fig. 2 is a schematic diagram showing the action of the chain as it enters a pulley. A phenomenon called the polygon effect occurs because the chain has a finite pitch l p . As a result, chordal vibration of the chain occurs. The amount of chordal vibration is defined as chordal displacement z . Fig. 3 shows a simple model of a chain entering a pulley. In this model, chordal displacement z is expressed by Eq. (1) . z = R ( 1 \u2212 cos \u03b8 ) \u03b8 = sin \u22121 l p (1) 2 R From Eq. (1) , it can be seen that chordal displacement z increases when the pulley applied radius R is small or when chain pitch l p is large. In addition, when the input rotation speed is N in (rpm), the noise of a chain-type CVT increases at frequency f 1 st , caused by pitch l p as shown in Eq. (2) [2] . f 1 st = 2 \u03c0R in l p N in 60 (2) To indirectly identify the behavior of a CVT chain that causes noise, the acceleration of the pulley was measured under conditions that cause the sound pressure at frequency f 1 st to increase [2] ", + " This property affects the behavior of the tension acting line, that is, chordal displacement z . This property can be changed by the cross-sectional curve of pin A \u03b6 p 1 . Fig. 8 shows the relationship between the movement amount S k +1 ( \u03b1k ) of the contact points between the pins with respect to the bending angle \u03b1k between the links. (a) shows S k +1 ( \u03b1k ) when the cross-sectional curve of pin A \u03b6 p 1 is involute. (b) shows S k +1 ( \u03b1k ) when the tension action line does not move when the pulley applied radius R is the minimum of 30 mm. As shown in Eq. (1) and Fig. 3 , the highest chordal displacement z is attained when the pulley applied radius R is the minimum value. The cross-sectional curve of pin A in (b) is a characteristic that sets chordal displacement z in this case to 0 mm. However, S k +1 ( \u03b1k ) in (b) is 0 mm when the bending angle \u03b1k is from 0 to 0.1 rad, and the radius of curvature is 0. As a result, the stress becomes infinite at these points when the pins contact each other. Therefore, it cannot be practically used a curve when the bending angle \u03b1 is from k 0 to 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure8-1.png", + "caption": "Figure 8. Oil sand units with 2 DOFs connected by spherical joints [10,24].", + "texts": [ + " KRD u14; t\u00f0 \u00de \u00bc cos h14 _w14 \u00fe _/14 x0 z14 \u00bc 0 (20) \u03c814 is the rotation angle about original z-axis; \u03b814 \u2013 rotation about the new x-axis; \u03d514\u2013 rotation about the new z-axis; x0 z14 \u2013 angular velocity about the z-axis of body- fixed coordinate system. x0 z14 can be assumed similar to translation motion in Equation (19) as x0 z14 \u00bc 0:12 t2 0:016 t3 ; t1 t t2 1:0 ; t[ t2 deg=s (21) Due to non-holonomic rotational constraint, the crawler track assembly still has five degrees of freedom (two translations and three rotations) to move freely during the turning motion of the crawler track assembly. A section of flexible oil sand terrain model made of 50 oil sand units (bodies 15\u201364 in Figure 3) connected by spherical joints is shown in Figure 8(a) from Frimpong and Li [10] and Frimpong et al. [24]. Each oil sand unit has four motion constraint equations D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 (two translational and two rotational) at its body fixed to centroidal coordinate system that allow only 2 DOFs. This 2 DOFs oil sand unit can be represented as a simple spring-mass-damper system as shown in Figure 8(b) from Frimpong and Li [10] and Frimpong et al. [24]. However, when oil sand units with 2 DOFs are connected by spherical joints, redundant constraints equations are introduced in the oil sand model. To eliminate the redundancies, the combination of spherical and primitive in-plane joints is used to connect the oil sand units (Figure 9). In addition, two translation motion constraints on the oil sand units 1\u201349 (Parts 15\u201363) are also removed. The constraint equations for spherical joints connecting two adjacent oil sand units shown in Figure 9 are similar in form to Equations (13) and (14) and hence not shown here" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002599_s1068366616040024-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002599_s1068366616040024-Figure1-1.png", + "caption": "Fig. 1. Scheme of the damper with the compressible oil film and porous cage: (1) oil film, (2) external ring of the roller bearing (nonrotating), (3) support springs, (4) axle of the damper with the pressed film, (5) axle, and (6) porous cage.", + "texts": [ + " This paper is devoted to the development of the model of damper gain with the compressible oil layer and porous insert with the variable permittivity under the conditions of the incomplete filling of the clearance with the lubricant material assuming the feed of the lubricant material at radial and circumferential directions. The aim of this work is to generate the specified calculation model of the damper with porous sintered insert and compressible oil layer, which provides the assumption of the simultaneous effect of the set of variable structural and operational factors on the gain. STATEMENT OF PROBLEM The scheme of the damper with a compressible oil film and a porous cage is given in Fig. 1. The equation of motion of rotor in directions r and t (Fig. 2) for nonstationary motion of the axle center can be represented as follows: (1) (2) ( ) ( )[ ] ( ) ( ) 22 2 2 sin cos cos , r Y Y X X dd em e F W K Y dtdt K X u t \u23a1 \u23a4\u03d5\u2212 = \u2212 \u2212 + \u03b4 \u03d5\u23a2 \u23a5 \u23a3 \u23a6 \u2212 + \u03b4 \u03d5 + \u03c9 \u03d5 \u2212 \u03c9 ( ) ( ) ( )[ ] ( ) ( ) 2 2 2 2 cos sin sin . t Y Y X X d ddem e dt dtdt F W K Y K X u t \u23a1 \u23a4\u03d5 \u03d5+\u23a2 \u23a5 \u23a3 \u23a6 = \u2212 \u2212 + \u03b4 \u03d5 + + \u03b4 \u03d5 \u2212 \u03c9 \u03d5 \u2212 \u03c9 396 JOURNAL OF FRICTION AND WEAR Vol. 37 No. 4 2016 AKHVERDIEV et al. Assume the working load W is stationary and oriented according to the scheme in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000872_iweca.2014.6845682-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000872_iweca.2014.6845682-Figure8-1.png", + "caption": "Figure 8. Dynamic stresses of longitudinal direction.", + "texts": [ + " An example of longitudinal direction strain in position one is illustrated in Fig. 6. The comparison between harmonic analysis and experiment is listed in table II. It is acceptable since the maximum error between them does not exceed 10 %. The errors mainly result from the instruments and the numerical algorithm. Eventually, dynamic stress amplitudes of the two positions obtained by the harmonic analysis are presented in Fig. 7. The tendency is similar to that in Fig. 5. Dynamic stresses of two directions for the entire workpiece are presented in Fig. 8 and 9. Conclusion is drawn that the dynamic responses differ from one position to another. As the workpiece is not regular, it is impracticable to calculate the dynamic stresses with the strain results obtained by the experiment using the theory of elasticity. A further study of the relationship between dynamic stress amplitudes of two positions and exciting forces are shown in Fig. 10. The linearity implies that the larger exciting force is given, the stronger response is to get. Altogether the FEA in VSR process does indeed provide significant results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002344_s00170-016-9132-0-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002344_s00170-016-9132-0-Figure2-1.png", + "caption": "Fig. 2 Car door trim with IS joints", + "texts": [ + " Fortunately, finite element method (FEM)-based approaches using well-defined material properties and numerical models can be excellent alternatives to physical experiments, producing reliable simulations of joining processes [2\u20135]. Results indicate that FEM is a powerful technique for predicting the performance of a joining technique. To increase the application potential of the joining technique, the development of a new staking process using infrared energy for an automotive door trim with polypropylene (PP) material (Fig. 2) is considered herein. Door trim is a common component in the automotive industry and is manufactured in large quantities. It is essential to have a reliable numerical model for conducting parametric studies in order to improve joint strength. Within a process chain, the outputs of preceding processes are considered to be inputs for succeeding steps. Moreover, we found that altering the process parameters, such as heating time, cooling time, and airflow rate, leads to variations in the mechanical behavior of the infrared staking joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000986_gt2014-26756-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000986_gt2014-26756-Figure1-1.png", + "caption": "FIGURE 1. SCHEMATIC OF BEARING CHAMBER FLOW REGIMES", + "texts": [ + " Recent research focuses on a possible numerical simulation of the flow inside a bearing chamber (for example Wang et al. [7] or Peduto et al. [8]) as well as on a better understanding of the relations and driving parameters through experiments. Performing experiments, Kurz et al. [9] identified two distinguished flow regimes and showed their influence on the film flow inside bearing chambers. The first regime occurred at lower shaft speeds, where the oil appears to be mainly driven by gravity. Figure 1(a) shows a simplified schematic of that regime. The second regime, which occurred at higher shaft speeds, was characterized by a more homogeneous and rotating oil film (Figure 1(c)). Kurz et al. [10] further discussed the bearing chamber regimes and experimentally determined a significant influence of the flow regime on the scavenge efficiency, a parameter which describes how much oil is scavenged. They also mentioned that the change from one regime to the other occurred at a specific shaft speed. Further research on the film thickness distribution in bearing chambers by Kurz et al. [11] revealed that a reduced chamber pressure increased the critical shaft speed at which the transition between the regimes was observed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure1-1.png", + "caption": "FIGURE 1. SCHEMATIC REPRESENTATION OF STRAIGHT VBD BRUSH SEAL.", + "texts": [ + " The stiffness of the seals is important to consider for assessment of rotor stability. An ideal location for introducing these seals is the end packing location on steam turbines. The seal provides a potential for reduced span in turbine end packing sections by replacing multiple labyrinth or interlocking teeth seals with single brush seal, reduced turbine end packing section leakage per axial length compared to labyrinth or interlocking seals, and improved sustained performance compared to labyrinth or interlocking teeth seals. The VBD brush seal (Figure 1) employs a front layer with thicker bristles to make the brush seal more stable under the turbulence of incoming flow. Although the idea of using thick bristles near the backplate to withstand pressure loading, as applied to the multi-layered brush seal (Patent # US5480165) is retained, the new feature in the VBD brush seal is the fine bristles sandwiched between a layer of thicker bristles and a layer of thickest bristles so that the fine bristles are protected from wear [4] and flow-induced vibration and high bending stress, while the fine bristles, as the core of the seal, do the sealing work and provide flexibility" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003288_1350650120936859-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003288_1350650120936859-Figure1-1.png", + "caption": "Figure 1. Geometry of offset-halves bearing.", + "texts": [ + " In the literature, studies dealing with micropolar lubricated offset bearings are limited. Chetti22 worked with micropolar lubricated offset-halves bearing and carried out static and dynamic analysis for a single input parameter, i.e., aspect ratio. Hence, it is important to study the effect of aspect ratio and offset factors on the stability of offset-halves journal bearings. The author makes an attempt in this direction in order to understand the relationship between stability and offset factor, stability and aspect ratio, etc. The geometry of offset journal bearing is shown in Figure 1. The center of each half is shifted by a distance XL from the bearing center along the split axis. XL is called the horizontal preset. The center of the upper half is on the right side while the center of the lower half is on the left side for counterclockwise direction of rotation. The centers will interchange for the clockwise direction of the journal. For a concentric shaft position, there are two reference clearances: a minor clearance and a major clearance. Physically, a micropolar model can represent fluids whose molecules can rotate independently of the fluid stream flow and its local vorticity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001388_isie.2014.6864934-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001388_isie.2014.6864934-Figure6-1.png", + "caption": "Fig. 6. Arms rotation in the operation of a disconnect switch", + "texts": [ + " The discharging process was tested under different operation modes, with the following results: \u2022 In case the sensor was programmed to communicate every 10s, the energy stored in the fully charged capacitors would allow operation during 2 hours and 30 minutes; \u2022 In case it was programmed to communicate every hour, operation would last for 10 hours; \u2022 In the standard factory programming of this smart sensor, temperature is hourly monitored, and communication occurs twice a day in normal circumstances, which corresponds to power autonomy of 15 hours. Fig. 4 shows three discharge curves, one for each operation mode. B. Optoelectronic Sensor The alignment between the two parts of a disconnect switch is measured by the optoelectronic sensor located in one part only (female), as illustrated in Fig. 5, the optical coupling being ensured by a reflective surface located in the opposite male part, as both parts rotate in the same plan Fig. 6. Thus, whenever both arms in a HV switch sensing device and the reflective surface are each other (alignment angle 0\u00ba) and only phototransistors are irradiated with infra-red light two parts become misaligned, the angle between them implies that the reflected infra-red beam irradiates other than the central ones, thus allowing the degree of misalignment to be determined. In the process of measurement the degree of infra-red LED is turned on and, after an allowed settling time of 50 ms, logic states at the outputs of all phototra evaluated in order to determine the angle between the two arms in a HV switch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003565_j.apm.2020.09.018-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003565_j.apm.2020.09.018-Figure3-1.png", + "caption": "Fig. 3. Schematic diagram of the eccentric structure.", + "texts": [ + " One rotation of the electromagnetic field corresponds to two tooth distances of the movable- tooth motion; consequently, a reduction ratio occurs, and a large output torque can be obtained. 3. Electromagnetic force and flexible-wheel deformation under eccentricity When an eccentricity occurs in electromagnetic harmonic movable tooth drive system, the initial air gap in the system is unevenly distributed. The center O of the center wheel is taken as the origin to establish the coordinate system, and the eccentricity is denoted by h . The structure of an eccentric drive system is shown in Fig. 3 . There, r is the radius of the flexible wheel, R is the radius of the center wheel, R \u2019 is the distance from point O to one point A on the flexible wheel, \u03b40 is the initial air gap without eccentricity ( \u03b40 = R - r ), \u03b8 is the angle-coordinate of point A , and q r is the dynamic electromagnetic force per unit length on the flexible ring. From Fig. 3 , it can be known { R \u2032 = \u221a x 2 + y 2 y = R \u2032 cos \u03b8 \u03b4 = R \u2212 R \u2032 (1) x 2 + (y \u2212 h ) 2 = r 2 (2) where x and y are the x-coordinate and y-coordinate of point A. From Eq. (2) , it is known that x 2 + y 2 = r 2 - h 2 + 2 hy . Considering the first and second equations of Eq. (1) , yields R \u2032 2 = r 2 \u2212 h 2 + 2 hR \u2032 cos \u03b8 . So R \u2032 = 2 h cos \u03b8 \u00b1 \u221a 4 h 2 c o s 2 \u03b8 \u2212 4 ( h 2 \u2212 r 2 ) 2 = h cos \u03b8 \u00b1 \u221a r 2 \u2212 h 2 (1 \u2212 co s 2 \u03b8 ) Taking a positive value, yields R \u2032 = h cos \u03b8 + \u221a r 2 \u2212 h 2 (1 \u2212 cos 2 \u03b8 ) (3) Substituting Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000962_j.proeng.2013.12.149-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000962_j.proeng.2013.12.149-Figure1-1.png", + "caption": "Fig. 1. (a) Journal bearing test rig diagram and (b) Pressure sensors attached to the bearing system.", + "texts": [ + "s), U is rotational speed (rps), L is bearing length (m), R is journal radius (m), c is clearance (m) and is eccentricity ratio. This equation reduces to the simplified expression for Petroff friction when the shaft and bush are concentric. There are three important assumptions in Eq. 1: (i) Bearing surfaces are smooth, (ii) The fluid is Newtonian and the flow is laminar, and (iii) Inertia force resulting from acceleration of the fluid and body forces are small compared with the surface force. The Journal Bearing test rig in Fig. 1(a) was used in this experiment. The loading arm is mounted to the bearing. The frictional force sensor is mounted on the spindle housing as shown in Fig. 1(b). When the loading arm presses the loading pin, the force sensor will record the friction values. A pneumatic bellow is used to apply the required load. The maximum speed of the journal test rig is 1000 rpm. The speed values used for testing were 500 and 800 rpm. The tests were conducted at different loads (10, 15 kN). Oil pressure supply was set at 0.2, 0.5 and 0.7 MPa. The groove position was varied at -450, -300, -150, 00, +150, +300 and +450 from the loading arm. Details of test bearing dimensions, lubricant properties and operating parameters are given in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000630_ilt-11-2011-0103-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000630_ilt-11-2011-0103-Figure4-1.png", + "caption": "Figure 4 Film thickness model and the reference system", + "texts": [ + " As cc is the smaller clearance of the journal housing, the two \u201cbarrier\u201d surfaces should reduce the lubricant side loss effectively also in nominal working conditions. In summary, the bearing (out of the barrier zones) is made up by three pads: two eccentric pads of the \u201celliptical\u201d part (with eccentricity e1, e2, respectively, and clearance c1 c2 cp) and a pocket pad (with eccentricity e3 0 and clearance c3 cc ). At this point, the eccentricities ei and clearances ci of the three pads (i 1, 2 and 3) are known. As shown in Figure 4, in the bearing reference system Oc, X, Y the locations of OJ (journal center) and Oi (pad i center of curvature) are determined by means of the Cartesian and polar coordinates eX, eY and ei, i, respectively (where 1 180\u00b0 and 2 0). Hence, by neglecting higher-order infinitesimals and misalignments, the film thickness h at the polar coordinate for pad i is given by: h ci (eX ei cos i)cos (eY ei sin i)sin (10) The deviation of an actual dimension of the bearing from the basic size will be denoted in the following by means of the prefix ", + " They are listed in Table I together with the corresponding driven dimensions, calculated by suitably introducing the driving dimensions in the equations reported in the previous paragraphs. Table II reports all of the data needed by the lubrication analyses, performed for LMC and MMC geometries, in nominal conditions. The lubricant is ISO VG 46 oil. The bearing is equipped with two feed grooves, opposite with reference to the load line, but is fed only through the groove upstream the active film region. It is characterized by means of the circumferential coordinate s of the center; the angular size s, as shown in Figure 4; and the axial length Ls. Their average thickness ts is used only to take into account of dissipation in the grooves by means of Wendt empirical correlation (Stefani and Rebora, 2009). The feed pressure ps is increased from 0.07 MPa (nominal value) to 0.1 MPa to assess the corresponding rise in the lubricant flow rate. All of the thermal models are used in conjunction with the mass-conserving cavitation algorithm (mc), except in the case of thermal Model 1, also used together with a quasi-static treatment of cavitation (qs)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000985_j.procir.2013.08.001-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000985_j.procir.2013.08.001-Figure1-1.png", + "caption": "Fig. 1. Typical layout of drive train components in a modern wind turbine, with a hub carrying three blades, a main low speed shaft, a step-up gearbox and a high speed connection shaft to the generator", + "texts": [ + " Until a few years ago deployment growth rates in Europe were around 25-30% annually and even nowadays similar growth rates are seen worldwide. Despite the recession in the EU over the last few years growth rates of around 10% are maintained in this part of the world. In the beginning of 2013 a total of 264GW was installed worldwide, and with an average capacity of 1MW, this means that more than 250,000 machines are currently in operation. Since the majority of these wind turbines are designed using a standard lay out using a low speed main shaft connected to the hub with the rotor blades, a gearbox and a generator, (Figure 1), it is evident that wind turbines are a large market for gear box manufacturers. Contrarily to ordinary industrial gearboxes wind turbine gearboxes are designed for a reversed mode of operation, since power needs to be transferred from the low speed shaft to the high speed shaft. Furthermore the ingoing rotational speeds are extremely low, typically in the range of 10-20 RPM where outgoing rotationally speeds usually comply with the range common to Megawatt scale electrical motor/generators, hence in the order of 750-1500 RPM", + " The same holds for yaw misalignment. This is a situation in which the wind turbine nacelle is not properly aligned with the wind direction, and this also gives yield to cyclic load fluctuations on the blades. In principle the control system tracks wind direction changes, so in a (10 minute) time averaged sense there is no yaw misalignment. But within a 10 minute average direction large scale turbulences also create wind direction changes of typically + 10-15 degrees, and the yaw system (yaw bearing and yaw motors, shown in Figure 1) is too slow to anticipate these wind directional changes. With wind turbine blades in blade length equal (or even larger) than the span of the largest commercial aircraft it might be wise to try to mitigate such blade load fluctuations. IPC or CPC pitching actions are a possibility to reduce such blade load fluctuations. This has been identified by a number of wind turbine manufacturers and few of them, a.o. Vestas, GE-Wind, Siemens and MHI (Mitsubishi) have implemented IPC or CPC in their products" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000344_gt2015-43161-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000344_gt2015-43161-Figure6-1.png", + "caption": "Fig. 6 Coordinate system of a floating bush bearing", + "texts": [ + " By using the following expression, the coordinate values ( , Z) in the orthogonal coordinate system were converted into the coordinate values ( , ) in an oblique coordinate system. sinZ (3) tan/Z (4) The nondimensional mass flow rates Q and Q in and axis directions are obtained as follows under the assumption that the airflow in the bearing clearance is isothermal and viscous and the flow is laminar. 2tan 1 sin 1 2 3 1 ZPH Z PPPHQ (5) 2tan 13 1 Z PPPHQ (6) In this paper, the nonlinear orbit method (NOM) was used to calculate the high-speed instability. In NOM, the orbit of the rotor is obtained by solving the Reynolds equation and the equation of motion simultaneously. Figure 6 shows the coordinate system for NOM. In this coordinate system, when the translational whirling of the rotor is taken into consideration, the bearing clearances in the land and groove regions are given as follows. Land region: cossin1 byybxxrH (7) Groove region: cH byybxxg cossin1 (8) In the time step n, the equilibrium between the mass flow rates flowing into and out of the control volume is given as follows. 1 ,, , 1 ,, , 43214321 2 n ji n jin ji n ji n jin ji nnnnnnnnn PP H HH PZ QQQQQQQQQ (9) The semi-implicit Crank\u2013Nicolson formulation was applied for solving Eq. (9). The following boundary conditions were assumed in the following calculations. 20 |P|P (10) 1|| 22 lZlZ PP (11) Assuming that the whirling motions mode of the rotor and bearing bush is only in translational mode, we can obtain the following equations of motion in the x and y directions using the numerical model indicated in Fig. 6. For a rotor: 4 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use d d BKdZdP l r d d M by wbyw r l r l by b 2 2 2 0 2 2 cos 2 (15) To obtain the orbits of a rotor supported by the proposed flexibly supported aerodynamic journal bearings, the Reynolds equation and the equation of motion of the rotor must be solved simultaneously using the following calculation procedure. (1) The pressure distribution for the concentric condition of the rotor is solved numerically for certain values of dimensionless mass parameter M* bearing number and then the initial positions of the rotor and the bearing bush at time t = 0 are given at ex = 0, ey = \u20130" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002470_iccsce.2014.7072728-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002470_iccsce.2014.7072728-Figure2-1.png", + "caption": "Fig. 2: Radial magnetic force acts on the stator. Surface vibration thus leads to emission of electromagnetic acoustic noise.", + "texts": [], + "surrounding_texts": [ + "Keywords\u2014Electromagnetic acoustic noise, induction motor, inverter, space vector pulse width modulation, spectral test.\nI. INTRODUCTION Due to its well-known ruggedness and simple construction, induction motor has found wide application in industries. Increasing industrial performance requirements has witnessed dramatic improvements in speed and torque control of the induction motor. The access to superior computation using microcontrollers and advent of modern solid state switches such as Insulated Gate Bipolar Transistor (IGBT) has led to development of what is the modern day inverter. Unlike the inefficient method such as pole changing and line voltage change, modern induction motor control method are based on variable frequency control. Variable frequency control is based on the concept of using power electronics to change the level of output voltage and supply frequency. This is achieved by using pulse width modulation (PWM) technique. Classification of various methods in variable frequency control is shown in flowchart in Figure 1 [1]. Thanks to its better dynamic properties, vector control method has rapidly displaced scalar control methods in many advanced application. Within vector control method, Direct Torque Control (DTC) and Field Oriented Control (FOC) inverter are widely available in the market. Both this method had led to improved efficiency and precise control of motor\u2019s speed and torque. Some of the important applications in which the DTC and FOC has is CNC machine, process automation and HVAC. Literature survey indicates that the both of this\nmethod are still constantly being investigated and improved up to today.\nWhile the developments in modern motor control methods were very impressive, the use of power electronics for switching has led to the emission of electromagnetic acoustic noise. Non-sinusoidal output voltage and current from inverter which has higher harmonic content due to switching has led to high frequency electromagnetic acoustic noise [2]. In the field of motor control, the acoustic noise may be of secondary importance compared to robustness of control methods. Nevertheless, the electromagnetic noise which occurs between the ranges of 2-10 kHz, can be very sensitive to human psychoacoustics and can thus lead to discomfort for human operator of the induction motor [3]. Radiated electromagnetic noise from a motor driven by an inverter is due to the interaction between radial magnetic flux in the air gap and motor\u2019s stator. The radial magnetic flux induces a force, Fm that tends to attract the stator.\n2\nWhere b flux density in the air gap \u00b50 vacuum permeability (4\u03c0 \u00d710-7 H/m) This magnetic force acts on the stator by deforming it and hence creates surface vibrations.\n(1)\n978-1-4799-5686-9/14/$31.00 \u00a92014 IEEE 267", + "The magnetic flux density in the air gap consists of fundamental and a lot of harmonics generated by spatial distribution of coils, air gap thickness, eccentricity of rotor and current harmonics in power supply. Thus, in order to operate an induction motor silently, the flux density harmonics needs to be reduced [4].\nIn this paper, an inverter with Space Vector PWM (SVPWM) output driving 0.75kW induction motor is investigated. The relationship between electromagnetic excitation, surface vibration response and radiated electromagnetic acoustic noise is examined. Methodology to investigate the relationship between the electromagnetic excitation and resulting radiated acoustic noise is presented. Effect of varying the motor speed on overall sound pressure level is also investigated. Fast Fourier Transform (FFT) was implemented on the measurements and the resulting spectra indicate a strong correlation between electromagnetic excitation and radiated noise. The observed relationship is then explained from Source-Path-Receiver model. Strategies to reduce electromagnetic noise are also briefly presented.\nII. EXPERIMENTAL METHODS\nThe specifications of induction motor and inverter used is shown in Table I and II.\nThe SVPWM waveform can be traced by measuring the line-to-line voltage on motor terminals [5]. From the waveform trace, frequency components of the electromagnetic excitation are obtained. Picotech TA 041 Active Differential oscilloscope probe was used in this case. The differential probe leads were connected to two of the three motor terminals (U, V&W). The probe\u2019s BNC end was connected to Agilent DSO-X 3034A oscilloscope as shown in Figure 3 and 4. Using the built-in function in the oscilloscope, the FFT was computed as shown in Figure 5. FFT was obtained for speeds from 0 to 1500 rpm with increment of 250 rpm." + ] + }, + { + "image_filename": "designv11_34_0002333_978-3-319-41468-3_6-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002333_978-3-319-41468-3_6-Figure4-1.png", + "caption": "Fig. 4 The orientation of samples in building chamber and final fabricated specimens: A\u2014 inclined samples with machining allowances, B\u2014horizontal as-built samples, C\u2014support structure", + "texts": [ + " The total energy density of the laser beam changes with the different hatch spacing [32] so some modifications in the microstructure and mechanical properties of material can be expected. Besides hatch spacing the energy density (E) depends on other parameters of laser sintering technology: laser power, scan velocity of laser beam and of layer thickness of powder bed: E \u00bc P v hst \u00f01\u00de where E (J mm3) is density of energy, P (W) means the power of the laser beam, v (mm/s) is the scan velocity, hs (mm) is the hatch spacing, and t (mm) is the thickness of the monolayer. The second key factor in the fabrication process was the angle orientation of samples in the building chamber\u2014Fig. 4. Three positions were set: horizontal \u03b1 = 0\u00b0, inclined \u03b1 = 45\u00b0 and vertical \u03b1 = 90\u00b0. Six specimens were built without machining allowances (as-built) and they were oriented only in the horizontal position (three with standard hatch spacing, and three with reduced hatch spacing). The settings of process parameters and the value of energy densities are in the Table 3. An overview of the fabricated specimens can be found in Table 4. All specimens were CNC machined with the turning centre SPN 12 and the Sinumerik 840D control system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002481_urai.2016.7625728-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002481_urai.2016.7625728-Figure2-1.png", + "caption": "Fig. 2 The implementation of a quadrotor with a GPS module and ZigBee.", + "texts": [], + "surrounding_texts": [ + "The technique is simple to use for the formation control in this study. Four drones will be moving the same for the same control signal from the laptop. And each of the drones share the position information of the GPS mounted on the drone in a mesh network of ZigBee. Each of the drones calculates how much to move through a shared position information. Before take-off, each of the drones calculates the distance between each drone from own position information and position information of the other drones. And each of the drones maintains a calculated distance during the flight. The initial position of the drone will be placed in the desired formation. After, each of the drones flies according to the control signals coming from the laptop. Fig. 1 shows that the control signal transmission. A communication module connected to a laptop can be seen that sending a control signal to each of the drones. 3. Implementation 3.1 Quadrotor specification The flight control unit (FCU) has used to implement is the Pixhawk of the 3 D Robotics, which is an outstanding company in this field. By modifying an open source offering the 3D Robotics, we used for FCU. And the other part of the drone, frames, brushless direct current (BLD C) motors, motor drivers and batteries, were purchase . d separately. Each drone was produced in the all same specI- fication. The size of the drone used in the experiments is 500mm. Also the size of a propeller is I I-inch. In general case, it is difficult to fix the drone in the air. The drone is flowing because of less friction in the air. Due to various reasons, these problems appear. Among them, the biggest reason is the location of the center of gravity of the drones. To position the center of gravity of the drones in the center is very difficult. So, in order to prevent that the drones flow, we correct the position by using GPS. As mentioned earlier, however, GPS has a positioning error of about several meters. For this reason, to reduce the positioning error, we used an optical camera with GPS. Optical camera knows own moving direction for finding the pattern of the floor. Position control mode using the optical camera and GPS so that makes drones only move for control signals. Typical current drones are only one drone operations thorugh a single transmitter. In order for the formation co- ntrol of 4 drones in the same direction at the same time, it must be connected to the four drones in one transmitter. For this implementation, we communicate I-to-4 communication using ZigBee. At this time, baud rate of ZigBee communication use 115600 bit per second and distance of laptop is within a radius of 200 meters. When master ZigBee connected to the laptop broadcasts the control signal to others, four slave Zig Bee connected to each of the drones catch the control signal and move to match the control signal. In order to control through a laptop for each of the drones, we should know exactly what the signal from the transmitter to send quadrotor. By analyzing the code in an open source offering in the company, you can see what the signal is going through the transmitter. Through this method we can know some important signals, which are arming signal, throttle, roll, pitch and yaw. After confirmation of these signals, it created a console program that can continue to send some signal previously confirmed on the laptop. And through the modified open source only the signal, which is not the transmitter signal, coming through the ZigBee connected to the drone is to make the movable quadrotor. 3.3 The additional equipment in system Fig. 3 is a ground control system provided by pixhawk.org installed on the laptop. Using this software program, we can monitor the value of each variable and the current state of the drones. In addition, the GPS position imformation is also displayed on the program. As previously mentioned in this study, the GPS module was used to estimate the position of a quadrotor in outdoor environment. And an optical camera was used to reduce the effect of a positioning error of the GPS module. Fig. 4 is a system model showing the entire system. It shows the correlation of the respective device." + ] + }, + { + "image_filename": "designv11_34_0003915_j.triboint.2021.106876-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003915_j.triboint.2021.106876-Figure1-1.png", + "caption": "Fig. 1. Test rig.", + "texts": [ + " The lubricant level and viscosity, rotational speed and rolling bearings geometrical specifications will be also considered as variables of the equation. To conclude, the sensitivity of the developed equation to parameters such as, speed, load, lubricant level and viscosity is analysed and compared with the SKF [13] and Witte [15] models. This is done in order to assure that there is an agreement between the developed and a current state of the art rolling bearing torque loss models. The test rig, shown in Fig. 1, is composed of a supporting structure (1) made of aluminium profiles, an AC motor (2) with speed control by frequency variation, a timing pulley (3) mounted on the motor output shaft and a second timing pulley (4) mounted on the upper shaft and a timing belt (5). The upper shaft (6) is supported by two ball bearings (7, 8) and it is connected to the test axle box (12) by two spring couplings (9, 11) and a torque transducer (10). The spring coupling (11) is a limiting torque coupling. The upper axle line, including the shaft (6), the rolling bearings (7, 8), the torque transducer (10) and the axle box (12) have been aligned using high precision laser technology, in order to Nomenclature \u03b1,\u03b2,\u03b8 equation exponents (\u2212 ) \u03b7 dynamic viscosity (N s m\u2212 2) \u03bc lubricant viscousity (cp) \u03bcbl boundary film coefficient of friction (-) \u03c9 angular velocity (rad s\u2212 1) \u03c90 reference angular speed (rad s\u2212 1) \u03c1oil oil density (kg m\u2212 3) \u03c5 kinematic viscosity (mm2 s\u2212 1) alub lubricant level from the shaft center (m) Cm dimensionless drag torque (\u2212 ) Cf m dimensionless constant (\u2212 ) C1,2 equation constants (\u2212 ) dm rolling bearing mean diameter (mm) dsh shaft diameter (m) Fa axial load (N) F* a axial load (lb) Fr Froude number (\u2212 ) G bearing geometric factor (\u2212 ) g gravitational acceleration (m s\u2212 2) h oil immersion depth (m) M0 no-load dependent torque loss (Nmm) M1 load dependent torque loss (Nmm) Mdrag frictional moment of drag losses (Nmm) Mhydrodynamic rolling bearing hydrodynamic torque loss (Nm) Mrr rolling frictional moment (Nmm) Mseal frictional moment of the seals (Nmm) Msi sliding frictional moment (Nmm) Mst rolling bearing starting torque (Nm) Mtransient rolling bearing transient torque loss (Nm) MWitte torque loss according to Witte (lb-in) n rotational speed (rpm) Rm rolling bearing mean radius (m) ro gear outer radius (m) Ref Reynolds number - Fossier model (\u2212 ) Rer Reynolds number - Rolling bearing friction torque loss model (\u2212 ) S1,S2 sliding friction moment (\u2212 ) Sm immersed surface area of the pinion (m2) T lubricant temperature (degrees) T1,2 rolling bearing width (m) TP" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.32-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.32-1.png", + "caption": "Figure 9.32 Pose of motion singularity with s2 = s4 tan \ud835\udf035", + "texts": [ + " The same equations also lead to the following task space compatibility equation, i.e. Eq. (9.546), which relates w3 to the angular velocity components \ud835\udf14\u2217 2 and \ud835\udf14\u2217 3. ?\u0307?6 = \u2212\ud835\udf14\u2217 3\u2215c\ud835\udf035 (9.544) ?\u0307?3 = \ud835\udf14\u2217 2 \u2212 ?\u0307?6s\ud835\udf035 = (\ud835\udf14\u2217 2c\ud835\udf035 + \ud835\udf14\u2217 3s\ud835\udf035)\u2215c\ud835\udf035 (9.545) w3 = \ud835\udf0e\u20323s4(\ud835\udf14\u2217 2c\ud835\udf035 + \ud835\udf14\u2217 3s\ud835\udf035)\u2215c\ud835\udf035 (9.546) (b) Second Version of the Motion Singularity This version occurs if s4s\ud835\udf035 = s2c\ud835\udf035 or s2 = s4 tan \ud835\udf035, while c\ud835\udf033 \u2260 0 and c\ud835\udf035 \u2260 0. In this version, the tip point P becomes coincident with the base frame origin O. This pose of singularity is illustrated in Figure 9.32 with a somewhat exaggerated appearance. It is exaggerated because actually it is hardly possible to bring P very close to O due to the physical shapes of the relevant links and joints. If the manipulator assumes the pose of this singularity with s4s\ud835\udf035 = s2c\ud835\udf035 and c\ud835\udf035 \u2260 0, Eq. (9.525) implies that the task-space compatibility condition becomes s4\ud835\udf14 \u2217 2 + s2\ud835\udf14 \u2217 3 = w3s\ud835\udf033 \u2212 (w1c\ud835\udf031 + w2s\ud835\udf031)c\ud835\udf033 (9.547) (c) Third Version of the Motion Singularity This version occurs if s4s\ud835\udf035 \u2260 s2c\ud835\udf035, while c\ud835\udf033 = c\ud835\udf035 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003419_aim43001.2020.9158833-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003419_aim43001.2020.9158833-Figure1-1.png", + "caption": "Fig. 1: Wheel-legged robot in Gazebo simulator.", + "texts": [ + " For this purpose, we start with studying the implications of the system modeling errors, external disturbances, and sensor noise on the motion stability of an off-the-shelf wheellegged robot called Igor from Hebi Robotics [15]. We model the robot in the ROS-Gazebo environment and examine the effects of not only translational push disturbances but also rotational push disturbances on the robot, which are not well reported in the previous works. Before designing the controller, we present the system dynamic model of the wheel-legged robot (Fig. 1). In our dynamic modeling of the robot, we assume a simplified inverted pendulum, where the body of the robot acts as a point mass. The origins of the robot-fixed frame \u03a3R and body-fixed frame \u03a3C are coincident and placed at the midpoint of the wheel axle. For pitch rotation \u03b2 and wheel torques, the forward direction of the robot is considered positive; whereas for yaw rotation \u03b1, counterclockwise motion is assumed positive. The three-dimensional model of the robot is represented in Fig.2, whereas Table I shows all the modeling parameters of the system", + "86 \u22121.66 \u221217.87 ] . is acquired using lqr command in MATLAB (Mathworks Inc., Natick, MA, USA) with system parameters listed in Table II. As the robot steers its heading using a differential drive mechanism, we note that the corresponding gain values for \u03b1 and \u03b1\u0307 states in Klqr have opposite signs. With the purpose of simulating our wheeled biped robot in the Gazebo simulator [21], we modeled our robot in URDF (Universal Robotic Description Format) using the dimensions of the Igor robot as shown in Fig. 1. The URDF model includes all the physical constraints of the real robot, such as joint limits, static friction, and damping coefficients, and the actuator\u2019s torque-speed characteristics. Gazebo is an open-source simulator that uses ODE [22] as its physics engine. ROS (Robot Operating System) is used to implement the control algorithm in C++, thus controlling the robot in Gazebo. The ROS-Gazebo simulation runs at a frequency of 1 kHz with a motion controller steering the wheeled biped robot in the desired direction while keeping the robot in an upright position at the same time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003243_j.ast.2020.105972-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003243_j.ast.2020.105972-Figure1-1.png", + "caption": "Fig. 1. Structure of the FTRV with single moving mass.", + "texts": [ + " In Section 3, a prescribed performance controller design is presented, including the auxiliary system, the command filter, the performance function and the error transformation function. In order to improve the system robustness, the disturbance observer (DOB) is proposed. The finite-time convergence is proved via the Lyapunov stability theorem. In Section 4, numerical simulations are performed on the FTRV, testing the performance of the controller. Finally, the conclusion is provided in Section 5. Considering one typical FTRV actuated by the single moving mass shown in Fig. 1 [30,31]. The FTRV mainly consists of two parts, the body B and the moving mass S , in which S moves along the trail parallel to zBaxis. The masses of the body and the moving mass are mB and mS . The mass of the FTRV is mT = mB + mS . The position and the velocity of the moving mass along zB-axis are denoted as zA and z\u0307A . Remark 1. The FTRV referred in this paper is controlled by the single moving mass merely. The centroid of the FTRV is time-varying during the reentry phase, providing the extra aerodynamic moment", + " where, \u03c9 = [ \u03c9x \u03c9y \u03c9z ]T is the angular rate of the FTRV; J 3\u00d73 is the rotation inertia matrix and can be described as J = \u2223\u2223\u2223\u2223\u2223\u2223 J11 J12 J13 J21 J22 J23 J31 J32 J33 \u2223\u2223\u2223\u2223\u2223\u2223 = \u2223\u2223\u2223\u2223\u2223\u2223\u2223 J110 + \u03bcz2 A \u2212 J120 \u2212\u03bcxoffset zA \u2212 J120 J220 + \u03bc(z2 A + x2 offset) 0 \u2212\u03bcxoffset zA 0 J330 + \u03bcx2 offset \u2223\u2223\u2223\u2223\u2223\u2223\u2223 where, J110, J120, J220 and J330 are constant parameters; zA , z\u0307A are the placement and the velocity of the moving mass, respectively; xoffset is the distance between the rail and the centroid of the body, which can be also shown as Fig. 1; mass ratio \u03bc is defined as \u03bc = mSmB/(mS + mB). Assumption. As the mass of the moving mass is far smaller than that of the reentry vehicle, the momentum moment caused by the velocity and acceleration of the moving mass is ignored. According to the theorem of momentum moment and based on the aforementioned assumption, the dynamics model of the FTRV is written as dH dt = \u2202 H \u2202t + \u03c9 \u00d7 H = M0 + Mextra (3) where, M0 denotes the initial aerodynamic moment. From (1), (2) and (3), it can be obtained\u23a1 \u23a3 \u03c9\u0307x \u03c9\u0307y \u03c9\u0307z \u23a4 \u23a6 = J \u22121( \u23a1 \u23a3 Mx0 + Mextrax M y0 + Mextray Mz0 + Mextraz \u23a4 \u23a6 \u2212 J\u0307\u03c9 \u2212 \u03c9 \u00d7 ( J\u03c9)) (4) where, J i j (i, j = 1, 2, 3) are the elements of matrix J ; D , L and Z are the aerodynamic forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001636_1.4032400-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001636_1.4032400-Figure2-1.png", + "caption": "Fig. 2 The forming of a GIE unit curve (rack method)", + "texts": [ + "\u201d Figure 1 shows the relative positions of the teeth-skipped worm and the gear to be measured during inspection. Two highresolution encoders are mounted on the axes of the worm and the gear, respectively. The GIE curves can be acquired by the data from the encoders. In order to investigate the meshing process between the teethskipped worm and the gear, it is assumed that there is only one tooth on the worm and only one tooth on the gear; thus, the full process of meshing is illustrated as shown in Fig. 2. As shown in lower part of Fig. 2, the complete quasi-static TE curve, including the approach stage, the involute stage, and the recess stage, is defined as the GIE unit curve in GIE technology. The arc EA2 _ , line segment A2A1 , and line segment A1F compose the complete trace of contact point in meshing process. The arc EA2 _ is the approach stage corresponding to the increasing segment of the GIE unit curve; the line segment A2A1 is the involute stage corresponding to the horizontal segment of the GIE unit curve; and the line segment A1F is the recession stage corresponding to the declining segment of the GIE unit curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000714_rose.2014.6952987-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000714_rose.2014.6952987-Figure5-1.png", + "caption": "Fig. 5: Orientation adaption of the gripper around the z-tool axis", + "texts": [ + " While the object is moving only its position will be estimated as the orientation is changing steadily and it is not necessary to follow the object without grasping. Once the conveyor stops the orientation of the object will be estimated as described above. To grasp the circular object it is not necessary to adapt the orientation. However, for the square and rectangular shaped objects grasping will be performed on two opposite sides. Otherwise, the triangle will be grasped on a side and its opposite corner. Fig. 5 illustrates how the orientation of the gripper will be changed according to the orientation and shape of the target object. V. GRASPING DURING CONVEYOR STOPPAGE The robot gripper has already the same position as the object in the y-z plane. The next step will be to move along the x-axis (depth information from Kinect)) towards the object (Fig. 1) and adapt the orientation of the gripper at the same time. To pull out the object successfully the force/torque control will now take over the control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003673_tmag.2020.3035280-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003673_tmag.2020.3035280-Figure7-1.png", + "caption": "Fig. 7. Mesh of the FEM model.", + "texts": [ + " However, once the parameterized FEM and LMC models for a specific type of electrical machine are established, they are applicable for a series of electrical machines of the same type by changing the parameters. Therefore, the creation time of FEM and LMC models can be neglected, while the efficiency of the FEM and LMC models mainly depends on the meshing and solution time. As shown in Table II, the speed advantage of the LMC models is prominent. The FEM model is established using ANSYS Maxwell software, and it is meshed automatically by the software to avoid overly dense mesh, as shown in Fig. 7. The solution speed of FLMC and SLMC models are approximately 14 times faster than the FEM, while that of the TLMC model is 6.5 times faster than the FEM. Besides, the meshing process of the FEM model is also time-consuming, which is not required in the LMC models. Although the solution speed of the TLMC model is slower compared to the FLMC and TLMC models, it will be shown later that the TLMC model exhibits higher accuracy. MODEL VALIDATION In this section, the electromagnetic performances of a 12s/10r DA-FSPM machine are predicted by the proposed LMC models and compared with the 2-D FEM prediction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002894_j.cja.2019.11.014-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002894_j.cja.2019.11.014-Figure4-1.png", + "caption": "Fig. 4 Schematic diagram of the typical engine structure.", + "texts": [ + " Section 2 simplifies the bearing-casing coupling system as a two-degree-of-freedom system, and proposes a double integral quantitative method for evaluating the influence of uncertain bearing stiffness and damping on the response amplitude and phase. In this section, a typical engine bearing-casing system is analyzed as a typical application, and the impact of uncertain bearing stiffness and damping on the response of measuring points is analyzed. itative evaluation of influence on thin-walled casing response caused by bearing A typical gas turbine engine structure is shown in Fig. 4. The engine has six bearings, of which No. 1, No. 2, No. 3 and No. 6 are low-pressure rotor bearings, and No. 4 and No. 5 are highpressure rotor bearings. According to the actual installation status of the engine, the constraint conditions are applied to the casing model. For the bearing-casing coupling model of the engine, the casing is modelled by 3-D solid finite element and the bearing are modelled by a kind of 2-D finite element as Ref. 27. The harmonic analysis is carried out by applying a constant rotating displacement excitation at the position of the bearing No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003421_aim43001.2020.9158930-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003421_aim43001.2020.9158930-Figure3-1.png", + "caption": "Fig. 3. Grasp pose that could reorient the object when the gripper grasps the object.", + "texts": [ + " At each orientation, GPD pushes the gripper forward from a collision-free configuration until the fingers of the gripper touch the input point cloud. These methods can generate grasp pose candidates for arbitrary types of objects and successfully perform pick and place tasks with cluttered objects. Since grasping in pick and place tasks and bin picking tasks has already been well developed, we focus on grasping conditions in assembly tasks. First, the change of the object pose should be minimized when the robot grasps the object because it can cause uncertainties, as illustrated in Fig. 3. Previously existing methods that consider only the approach direction face this problem. In addition, assembly of furniture typically requires a great amount of force and most furniture parts are plane-rich. Thus, we take advantage of the antipodal-based methods and approach-based methods to increase the contact area and consider using the grippers with the palm. In this section, we explain the generation process of grasp pose candidates. The candidates in the proposed method have a grasp position and orientation which describes the target pose of the gripper in SE(3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure2-1.png", + "caption": "Figure 2. The first experimental setup of the omnidirectional driving gear", + "texts": [ + " 1, when the spur gear on the left rotates to drive the omnidirectional driving gear in the direction of the X-axis, the spur gear on the right slides between the teeth of the omnidirectional gear with clearance for the backlash of the involute curve [1-3]. Conversely, if the spur gear on the right rotates to drive the omnidirectional gear in the direction of the Y-axis, the spur gear on the left slides between the teeth of the omnidirectional gear in the same way. By combining these spur gear motions, this omnidirectional driving gear can generate thrust in any direction. The first experimental setup with the planar omnidirectional driving gear is shown in Fig. 2. We also developed experimental setups with convex and Riichiro Tadakuma, Minoru Takagi, Shotaro Onishi, Gaku Matsui and Kyohei Ioka are with Yamagata University, Faculty of Engineering, 4-3-16 Jonan, Yonezawa city, Yamagata prefecture, 992-8510, Japan (tel: +81-238-26-3893; fax: +81-238-26-3205; e-mail: tadakuma@yz.yamagata-u.ac.jp, minoru.t@yz.yamagata-u.ac.jp, tsn39426@st.yamagata-u.ac.jp, txf38722@st.yamagata-u.ac.jp, tfh75982@st.yamagata-u.ac.jp,). Kenjiro Tadakuma is with Osaka University, Department of Mechanical Engineering, Room #406, Building-M4, 2-1 Yamadaoka, Suita city, Osaka, 565-0871, Japan (tel: +81-6-6879-7333; fax: +81-6-6879-4185; e-mail: tadakuma@mech" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003124_j.matpr.2020.04.239-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003124_j.matpr.2020.04.239-Figure5-1.png", + "caption": "Fig. 5. Hexa element meshed FEA model of drilling spindle.", + "texts": [ + " In the Ansys harmonic response tool, rotor dynamic results are incorporated to characterize the rotating structures. Three dimensional FEA model of the drilling spindle bearing system is generated and essential motion and load constraints are applied using rotor dynamic Ansys workbench tool. Tangential and axial movements are restricted at element 1 to assist simply support boundary condition by ignoring the shear effect; also rotation about y axis is achieved through radial motion of the spindle. The entire spindle bearing system is meshed using higher hexa mesh as shown in Fig. 5. Rotor dynamic and harmonic response analysis of spindle is carried for determining the natural frequencies, mode shapes, critical speeds, harmonic frequencies and peak amplitudes. Stability analysis is carried out by plotting Campbell diagram, The numerical analysis is performed at different speeds ranging from 0 to 36,000 rpm. Initially, natural frequencies of drilling spindle supported by angular contact bearings at locations 1 and 2 are obtained. Then the analysis is extended to drilling spindle supported by radial PMBs at 1 and 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002900_j.engstruct.2020.110218-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002900_j.engstruct.2020.110218-Figure11-1.png", + "caption": "Fig. 11. A simplified geometric model for clash angle calculation.", + "texts": [ + " All samples were manufactured without over-kerfing and so in most samples there was an evident second nonlinear response region which denotes the onset of clashing. The clash angle Au parameter was recorded at the commencement of this second nonlinear region and is summarised in Table 6. For samples where there was no obvious second nonlinear region, Au was recorded as the maximum fold angle. It is evident that thicker plates clash at smaller rotation angles. To explain this behaviour, a simplified geometric model is proposed which assumes the centre of rotation of one hinge leaf is at the edge of the adjacent leaf. With reference to Fig. 11a, if ak > t 2 p , a Case 1 clash will occur after the hinge leaf has rotated beyond 90\u00b0, with impact between the leaf corner and the top of the adjacent leaf. With reference to Fig. 11b, if ak < t 2 p , a Case 2 clash will occur before the hinge leaf has rotated 90\u00b0, with impact between the leaf side and the corner of the adjacent leaf. A critical case is encountered when =ak t 2 p , where impact will occur at 90\u00b0 in both locations simultaneously as shown in Fig. 11c. An expected clash angle can therefore be calculated as a function of plate thickness and measured kerf angle, using triangle geometry of Case 1 and Case 2, giving Eq. (6) as: = < a a 2tan for 2cos for t a k t a t k t 1 2 2 2 1 2 2 p k p k p p (6) Expected clash angles were calculated for all samples from measured kerf widths. Results are shown in Table 6, alongside measured clash angles Au. Case 1 is seen to govern for 1 mm plate; Case 2 for 3 mm plate, and either Case could govern for 2 mm plate, as the measured kerf widths are close to the critical value of t 2 p " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002153_8.8886-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002153_8.8886-Figure5-1.png", + "caption": "Fig 5b", + "texts": [], + "surrounding_texts": [ + "The Whitham theory predicting the far-flow field around a projectile is used to derive body shapes which produce extreme bow shock-wave pressure jump or \"boom,\" subject to constrain ing conditions regarding the drag due to the bow shock and fine ness ratio of the bodies. I t is found that the minimum drag body is also the minimum boom body. The body-volume effect and the effect of discontinuities in slope of the body meridian section on the boom intensity is investigated. As a general result of the investigation, it can be said that the boom of a projectile for given Mach number and flight altitude is primarily determined by its length and fineness ratio. The maximum variation in the boom intensity for pointed bodies with given length and fineness ratio is of the order of 10 per cent. The geometry of the bodies is thus found to play a minor role. (1) Introduction IN A PREVIOUS PAPER by Whitham,1 a method was developed for determining the flow field and the shock pattern of a supersonic projectile. The present paper investigates the stationary character of the in tegral giving the bow-shock pressure jump subject to different subsidiary conditions. According to reference 1, the wave drag of a super sonic projectile and the bow-shock pressure jump are two quantities which can be related to the same uni versal function, the F-iunction, describing the flow field. In the first part of the investigation that part of the wave drag which is due to the bow shock only is chosen as a subsidiary condition. A second subsidiary condition giving the maximum body thickness in terms of the F-iunction is also imposed on the problem. The body-volume effect on the bow shock and the effect of discontinuities in slope of the body meridian section are investigated in the second part of the paper. The variation of the boom in connection with volume for a certain family of bodies is studied. (2) General Considerations The formula predicting the pressure rise in the bow shock or the \"boom\" of a projectile is given in reference 1: i l / 2 5 = P ~ = h(y, Mm F(y) _ 0 dy (i) Received December 22, 1959. * Staff Scientist, Flight Sciences Laboratory. Now, Fluid Dynamics Research Group, Convair (San Diego). The author would like to thank Dr. Y. A. Yoler for his valuable criticism during the preparation of this paper. where p is the pressure immediately behind the bow shock, pm the pressure in the free stream, and h(y, Mm, r) is h(y, Mm, r) = 21 /4 7 ( 7 + iy1/2(Mm 1) 1/8 ^-3/4 (2) In Eq. (2), 7 is the ratio of the specific heats, Mm is the Mach number in the free stream, and r is the radial distance in the x, r coordinate system shown in Fig. 1. Under the assumption that the cross-sectional area of the projectile S(x) possesses continuous first and second derivatives S'(x) and S\"(x) and, further, that S'(x) and S\"(x) have a bounded variation, the function F(y) has the following form: F(y) \"\" 2TTJO S\"(x)dx y (3) For Eq. (3) to be valid, it is also assumed that the pro jectile has a pointed nose, that is, 5(0) = S'(0) = 0. The variable y appearing in the above formulas is de fined as the characteristic variable. On each char acteristic it has the value of x \u2014 ar at the point where the characteristic meets the body surface, a being the Straight Mach Line Fig. la F IG . 1. Continuous and discontinuous F(y)-functions related to actual body geometry. 113 D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 2, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .8 88 6 114 J O U R N A L O F T H E A E R O S P A C E S C I E N C E S \u2014 F E B R U A R Y , 1 9 6 1 value of (MJ \u2014 1)1 / 2 . The upper limit of the integral in Eq. (1) is defined as the first zero of the function F(y) apar t from y = 0 itself, t ha t is, 1 r Zir Jo \u2022y S\"(x)dx Vy0 - = 0 (4) Eq. (4) holds for those bodies t ha t have a con tinuously varying slope of the meridian section R(x) of the body in the interval 0 < x < yo + aR\u00b1 = X\\ (see Fig. l a ) . For this class of bodies it is seen t h a t the value of 3>0 is dependent upon S(x). However, in this paper we want to consider also another kind of a first zero of the F-function, namely, a zero which occurs when the F-iunction exhibits a jump a t \u2022 3/0 due to a discontinuity in slope of R(x) a t Xi = y0 + aR\u00b1 (see Fig. l b ) . For such bodies, Eq. (4) does not hold. In the following we want to consider both types of bodies. T h e physical situation together with typical ^-func tions is shown in Fig. 1. The shock wave star ts a t the pointed nose. All characteristics leaving the body surface up to the point Xi,Ri eventually coalesce with the shock wave in the far field. T h a t is, all disturb ances created by the body surface up to t ha t point feed into the shock wave and determine its pressure jump in the far field. The integral in Eq. (1) expresses this s ta tement as the area under the incurve up to the point y0. On the other hand, the drag of the body up to the point x\u00b1,Ri can be determined by integrating all pressure disturbances along the body surface itself. Another way of determining the drag is to take the integral of the entropy jump along the shock wave.1 Thus, using the Whi tham theory for the front shock only, the drag of the body up to the point Xi,R\u00b1 is given by D = irPa,U\u00ab \u201e F2(y) dy (5) where pm and Um are the free-stream density and ve locity, respectively. This formula, giving the dragdue to the bow shock, is different from the general drag formula for a body of revolution.2 This can be seen if Eq. (3) is substi tuted in Eq. (5) and one integra tion over y is performed. In the following, we shall use this formula to deduce relations between body shape, drag, and pressure jump of the bow shock in the far field.* (3) Bow Shock-Wave Drag and Boom Relations The first problem to be considered may be formulated as follows: Determine the cross-sectional area dis tribution S(x) in the interval 0 < x < X\\ of a pointed body of revolution which has a specified maximum section a t x = Xi, such t ha t this body gives an extreme value of the integral in Eq. (1) subject to the further condition tha t the drag of the bow shock given by * The author is indebted to Dr. G. B. Whitham for his sug gestion to use Eq. (5) as a representation for the drag of the body. Eq. (5) is prescribed. Thus, we want to find the sta tionary character of the integral T F(y) 0 dy (6) when the argument function F(y) is subject to the sub sidiary conditions: D TTp< >ua /\u2022yo = I F2(y) dy = a given constant G (7) J o (3/8)5(3/0) = a given constant C2 (8) A further condition which must be fulfilled by the Ffunction is F{y) > 0 for 0 < y < y0 (9) where the equalities may only occur simultaneously. In order to express the second subsidiary condition in terms of the F-i unction, Eq. (3), t reated as an Abel integral equation, can be written as d Cx F(t) dt dxJo Vx \u2014 / (10) The boundary conditions which can be specified for the pointed body are 5(0) = 0 and S'(0) = 0 (11) Performing two integrations of Eq. (10) and using the boundary conditions, Eq. (11) gives S(x) = I [^(0) *3/2 + \u00a3 F'(t)(x - tf'2 dt (12) I t is to be observed tha t the F-i unction is not necessarily zero for y = 0. The second subsidiary condition now takes the form g S(y0) = F(0)y0 3/2 + iF' (OCyo - 06/ dt = Q Introduce two Lagrange multipliers, X and ju; the in tegral to be studied is thus fyo pyo 1 = F(y) dy + A F\\y) dy + nF(0)y0 3/2 + J o J o M f0 F\\y)(y0 - y)3/2 dy (13) J o Suppose first t ha t Eq. (4) holds; namely, F(y) is con tinuous through 3/0 and F(y0) = 0. Then the variable y0 is dependent upon S(x). Therefore, when the varied function F(y, e) is introduced, e being a parameter, y& is changed and becomes a function of e, since Eq. (4) must also be satisfied for the varied function F(y, e). Thus, not only the integrands bu t also the limits of the integrals in Eq. (13) depend upon e. The total first variation of Eq. (13) can now be calculated and, in view of Eq. (4), we obtain 51 = I [8F + 2XF8F + nSF'(y0 - y)3/2} dy + 2 M 5y\u00b0 [I! F/(yKy\u00b0 ~ ^^ + my\u00b01/2] = \u00b0 (14) D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 2, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .8 88 6 S U P E R S O N I C B O O M O F A P R O J E C T I L E 115 where 81, 8F, etc., are defined by : dl(e)\\ 81 = 8F bF(y, e) First, let y0 be fixed and pu t the differential 8yQ = 0. After a partial integration of the third term in the first integral of Eq. (14), the Euler equation of the problem becomes F(y) 3 fx 4X (yo - y)1\" - _1_ 2X (15) Then, by letting y$ vary, we see t ha t the remaining ex pression in Eq. (14) is equal to zero: jufryo Jo F'(y)(yo-y)1/2dy F(0)y0 1/2 0 This expression cannot be zero, however, unless 8y0 itself is zero, since the expression within the brackets is the slope of the \u0302 -function a t yQ) and generally it is not zero. Therefore, yQ can be regarded simply as a con s tant in the calculation, which is to be determined a posteriori. Next, suppose t ha t Eq. (4) does not hold; t ha t is, suppose there is a jump in the 7^-curve a t y0. In this case y0 can immediately be regarded as a constant, and as a consequence Eq. (14) remains unchanged except for the last term which is automatically zero. The Lagrangian multipliers n and X are determined by means of Eqs. (7) and (8): A* = -(16/9;yo2)[C2X + (l/2)y0 3/2] (16) X = \u00b1( l /2)(3/ 0 ) 1 / 2 / (9C^o 2 - 8 G 2 ) 1 / 2 (17) Eqs. (15), (16), (17), and (9) now describe the solution completely. In the following, the constant G will be assigned the value G = (3/8)5Cvo) = ( 3 / 8 ) T T ^ 0 ! (18) Now, there exists a variety of solutions according to the value of G in Eq. (17) which satisfy the condition stated by Eq. (9) when the negative value of X is used. The lowest possible value which G can take, according t o E q . (17), is G = (l/8)(7T2i?o4/V) (19) Also, there exists an upper limit for the value of G which can be found by combining Eqs. (15), (16), and (17) with Eq. (9). This value of G is: G = (3/8)(x2i?0 4/;yo2) (20) The F-iunctions for the G values given by Eqs. (19) and (20) are, respectively, F{y) = 2 ^ 1 y_ 1/2 1/2-1 yJ J (21) (22) In general, the value of the integral T in Eq. (6) is found by substi tuting Eqs. (15) and (16) in Eq. (6): T = ( T T / 3 ) ( ^ O 2 / J O 1 / 2 ) ~ CVo/18X) (23) When X is an admissible negative quant i ty , it is seen t ha t T reaches its least value when X is as large as pos sible. The minimum value actually occurs when X tends to infinity, t h a t is, when G has its least value. The least value of G given by Eq. (19) is, however, the minimum wave drag due to the bow shock for a given maximum thickness of the body. This can readily be seen from Eqs. (17), (16), (15), and (14). According to these equations there exists only one La grange multiplier, say v, as X \u2014>* \u2014 \u00ab> with the value v = n/\\ = - ( 1 6 / 9 y 0 2 ) G The second variation of Eq. (13) is then: 82I o (8 F)2 dy>0 Thus, it can be concluded t ha t the minimum-boom body for given bow-shock wave drag and maximum body thickness is also the minimum drag body due to the bow shock for a given maximum body thickness. When X goes through the range of admissible values, the value of T increases monotonically and reaches its maximum value when G obtains the value given in Eq. (20) : Tmax = ( 7 r / 2 ) ( i V / V / 2 ) (24) The geometry of the noses which produce minimum and maximum boom or drag can now be found from Eq. (12). These bodies are, respectively, 1/2 R2(x) \u2022 = l i?o2 ( X yo 1 ) tanh\" 1/2- I (25) i?o2/ \u00a9 and \u2022 ^ (x)max \u2014 0 -\u0302 M) X \\ 1 / 2 yo/ (x 3\u0302o tanh\" - 3 x yj + \\ 1/2- I (26) Ro2g However, these bodies are to be extended up to the point Xi (see Fig. 1 or 2), where the body radius is R\u00b1. In order to be able to compare the results, it is there fore convenient to refer the bodies to the same body radius Ri a t x = X\\. Beyond the point xi, the body may be continued with, for example, a cylindrical por tion without affecting the bow shock. At the point Xi, there is then a discontinuity in slope of the body meridian section. This discontinuity causes the fluid to expand rapidly and gives rise to a recompression which appears as the tail shock in the far field, the drag of D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 2, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .8 88 6 116 J O U R N A L O F T H E A E R O S P A C E S C I E N C E S \u2014 F E B R U A R Y , 1 9 6 1 F I G . 2b. Geometry of maximum boom body and its .F-curve; Moo - V2 . which is given by -^rear shock ^Pea J F2(y) dy This drag, however, has no significance in the present problem because we are concerned with the front par t of the body exclusively. As a result of the discon t inui ty in slope a t xh the incurve exhibits a jump a t yo. T h e F-iunction then can take a manifold of dif ferent values a t the same place, indicating a Mach fan a t the shoulder of the body. Now for x = Xi, R2(Xl)n = R\u00b1 2 = Ro2f(~^ = Ro2fi and Ro2 = RiVh (27) Substi tut ing R0 2 from Eq. (27) in Eq. (25) gives the equation for the body radius in terms of xi and Rx. For a given Mach number, the proper value of y0 then can be found from the relation: yo Xi aRi I t is seen t ha t the functions / i and gi are functions of the product of a and the fineness ratio r, defined by The minimum and maximum values of T given by^Eq. (6) are as follows: mm - i/2 3/Aw 3 y o 1 / 2 3 / i W T = - ^ 2 V 7 2 2 & V3/0 Also, the drag of the bodies are: 1/2 T I / ^ L A ^3/2) (30) T'Xi Cnmin = C B M = D\u201e Dn {\\/2)p\u201eUJirRJ 7T2 R0* _ 4 y0 2i?i2 7T2 / * ! > 4 Vjv 3ir2 i?0 4 4 y o W ~ 3TT2 / X A r1' / l 2 1 (31) 4 \\ 3 V ft' For comparison, we choose the Mach number Mm = V 2 and r = 0.1. The values of Tmin, Tmax, CDmin, and CD a then are: Tmin = 1.002r2x1 3/2 CDm .B = 2.51 l r 2 Tmax = 1.200r2X!3/2 Cz>waiB = 4.802r2 As given by Eq. (1), the boom of the body is propor tional to the square root of T. The ratio of extreme boom intensities then is T 1/2 = f3)1/2 1/2 X / 25 5 4 \\ ( 1 - ar + T (ar)2 - - (ar)2 In - + . . . ) (32) \\ 16 8 ar J where the functions /1 and g\\ given by Eqs. (28) and (29) have been expanded for ar small compared to unity. For ar = 0.1, for example, \\-* max/ J- min) x.UzfD Thus, for slender bodies a t supersonic speed it can be stated tha t the maximum variation of the boom in tensity is of the order of 10 per cent. According to Eqs. (30) and (31), a general form of CD and T for any pointed body of revolution is CD = Ci(ar) r2 T = C2(ar) r2X?/2 D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 2, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .8 88 6 S U P E R S O N I C B O O M O F A P R O J E C T I L E 117 where Ci(ar) and C2(ar) are functions of ar. There fore, the boom of a projectile can be written in the form s = H[y,M~, (r/*i) , ar] CD 1/2 (33) showing tha t s is proportional to the square root of CD. In Fig. 2, the minimum and maximum boom body geometry is shown together with the ^-curves of the bodies for r = 0.1 and Mm = V 2 . (4) The Boom of Minimum Total Wave-Drag Bodies The results in the preceding section suggest tha t the minimum wave-drag bodies, according to the investi gations by Lighthill,3 Sears4 and Haack,5 also are minimum boom bodies. The minimum wave drag for a projectile with given maximum thickness and satisfying Eqs. (11) is, according to reference 3, CD = 4r2 The F-curve for this body is F(y) = (4/^)(i?1Vx1 3/2)(2\u00a3 - K) (34) where E and K are complete elliptic integrals with modulus k = (y/xx) 1/2 The F-curve given by Eq. (34) is very much like the F-curve given by Eq. (21) (see Fig. 2a). The value of the integral T for this body is* T = 0.983T!2:*:!372 (35) (5) Volume Effect on Bow-Shock Pressure Jump In this section we want to consider the volume of the projectile as another parameter of interest relative to the supersonic boom. The problem is to determine the distribution of a given total volume along a given length, such tha t the body thickness never exceeds a certain value, and the boom produced by this body is a minimum. In the following, we consider projectiles t ha t possess only one expansive zero of the F-iunction. For this class of bodies, the problem can be s tated as an isoperimetric variational problem of Eq. (6) with the volume up to point y$ given, since the bow shock is influenced only by the nose of the body up to yQ. How ever, this problem, when analyzed, gives only the trivial solution 5 = 0, which is in contradiction to the subsidiary condition imposed on the problem. There fore, there is no real physical solution to this problem. However, i t is interesting to s tudy the behavior of the relative minima for certain given families of bodies. Consider, for example, the family of pointed bodies defined by * I t is believed that this value of T is the lowest possible for any body of revolution with given maximum thickness and satisfying the Eq. (11). A proof of this statement is, however, very difficult to give and is omitted here. R(x) = Ri[l - (1 - x)n] n > 1, 0 < x < 1 R(x) = i?x x > 1 (36) Find the exponent n for which the integral T(n) = I F{y) dy J o ry0 = 2R1 2n I [{2n - 1)(1 - . T ) 2 ^ \" 1 ) - (n - 1)(1 - x)n-2Wy, -x dx (37) has its minimum value. The upper limit yQ(n) in Eq. (37) is generally to be determined from (\u0302:vo) = (2\u00bb J o y 0 ( l - x ) 2 ( w \" dx Vy0 \u2014 x / ; '0 V j o ~ x This equation is t rue for all n, which gives 0 1 \u2014 aRi, then y0 = 1 \u2014 aR\\ in Eq. (37). Thus, for a given Mach number we have two equations with two unknown variables n and 3/0 together with the inequality (39). This problem has been solved numerically for Mach number V 2. The results are shown in Fig. 3 with T(n) plotted against n. The volume of the bodies up to the point 3/0 is given by y* + --\u2014 (1 - yo)n+l -V(n) = irRx2 n + 1 ( l - .-yo)2B+1 2w + 1 3w + 1 X (M + 1)(2M + 1). This function together with yo(n) is also plotted in Fig. 3. The minimum of Eq. (37) occurs in the neighborhood of n = 2, for which T(2) = 1.025Ri2. As can be ex pected from the foregoing discussion, the minimum is very flat. The volume change around n = 2 is dV(n)/dn\\n=2~ O.lSTri?!2 which is considerable. The integral in Eq. (6) is thus insensitive to changes in the nose volume, being de- D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 2, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .8 88 6 118 J O U R N A L O F T H E A E R O S P A C E S C I E N C E S \u2014 F E B R U A R Y , 1 9 6 1 Fig 4a 0.02H F I G . 4. Fig 4b ^-curves for a cone-cylinder body with and without a spike; M^ = y/2. pendent primarily on the fineness ratio and body length. In this connection, i t is interesting to examine the Sears-Haack body producing minimum wave drag for a given body volume. The meridian section is given by R(x) = Ri[l - (1 - x ) 2 ] 3 / 4 0 < x < 2 where Rx = ( 8 F / 3 T T 2 ) 1 / 2 V being the volume of the body. I t is found t ha t T has the following value for this body: T = 0.988i?!2 This value is very close to t ha t given in Eq. (35). (6) Effect of a Discontinuity in Slope of the Meridian Section of the Nose on the Bow-Shock Pressure Jump In the preceding discussion we have considered bodies t ha t have a continuously varying cross-sectional area with continuous first and second derivatives in the interval 0 < y < yo. In this section, the effect of a discontinuity in S'(V) o n the bow-shock pressure jump is studied. For such bodies the T^-function cannot be calculated with Eq. (3). Instead, the /^-function must be evaluated with the Stieltjes integral1: xR(x)J 2T ^)=r by where h[(y \u2014 x)/aR{x)} is a function given in reference 1. In order to s tudy this question, consider a very slender cone (a spike) added to the bodies discussed in the foregoing section. At the spike-body junction, there is a sudden change in S'(x), giving rise to a second shock at tached to the body. In Figs. 4 and 5, respectively, the F-curves for the cases n = 1 and n = 3 are compared with the original /^-curves for the bodies. The ratio of the boom intensity of the original and the spiked bodies for the two cases is n = 1 - ^ ^ = 0.986 n = S ^ ^ = 1.021 ^sp. ogive Thus, a spike has very little effect on the boom and may lead to an increase in boom intensity, as in the case?z = 1. (7) C o n c l u s i o n The boom of a projectile for given Mach numbers and flight altitudes is determined primarily by its length and fineness ratio [see, e.g., Eqs. (30) and (35)]. The geometry of the projectile plays a minor role. For pointed bodies with given length and fineness ratio, the maximum variation in the boom intensity for a given Mach number and flight al t i tude is of the order of 10 per cent. The minimum wave-drag bodies thereby also are the minimum boom bodies. A general relation between the drag coefficient and the boom for any pointed body of revolution shows tha t the boom is proportional to the square root of the drag coefficient. (Continued on page 157) F I G . 5. F-curves for an ogive-cylinder body with and without a spike; M\u2122 = V 2 . D ow nl oa de d by U N IV E R SI T Y O F C A L IF O R N IA - D A V IS o n Fe br ua ry 1 2, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /8 .8 88 6 H I G H - S P E E D V I S C O U S C O R N E R F L O W 157 to their undisturbed two-dimensional values. The drag of the surface area disturbed by the merging is one half the drag due to two-dimensional shear over the same area. The total heat flux over this area is one fourth the two-dimensional heat flux over an equal area. (6) Values of momentum area, displacement area, and integration area are given. However, for the special conditions in which

stepped taper rod > taper rod. On the contrary, for the arrangement with L1 = 70 mm and L2 = 60 mm, the slope\nfirst slightly increases between stepped and stepped tapered which is followed by a decrease in the value between stepped tapered rod and tapered rod. No significant change in the slope is observed for rod having step in cross section. Contrary to this, in the case of taper rod with no step in their cross-section, shear stress is observed to be relatively lesser. Fig. 5 represents the behavior of cross-sectional deformation for the rod of 130 mm full length under the application of torque. For the arrangement of rod with L1 = 70 mm and L2 = 60 mm, the total deformation was observed to decrease while moving from uniform stepped rod to stepped tapered rod whereas it slightly increased while moving from stepped tapered rod to tapered rod. A continuously decreasing trend was observed for the rod having L1 = 75 mm and L2 = 55 mm with almost constant slope which was quite similar to the behavior of shear stress in the same rod. The sudden step in the uniform cross-section increases the deformation, whereas the step in the cross-section over the tapered rod was found to undergo comparatively lesser deformation. At the same time, it was also observed that the tapered rod and stepped tapered rod had almost identical amount of deformation. It may be concluded from the above observations that the step in the cross-section doesn\u2019t play a vital role in the case of tapered rod to provide resistance against deformation.\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 20,2020 at 23:37:24 UTC from IEEE Xplore. Restrictions apply.", + "Fig. 6 illustrates the strain energy distribution for the rod of 130 mm full length for the application of torque with the strain energy on the ordinate axis and the various types of components on the abscissa. From the fig. 6, a decreasing trend can be observed for both the types of arrangements of the rod. With the introduction of taper without step in the rod profile, strain energy was observed to become minimum, as opposed to the sudden step which significantly increased the strain energy. The decrease in the strain energy of the stepped tapered rod when compared to that of uniform rod was found to be steeper when L1 was selected as 75 mm, whereas it was observed to be decrement steeply in the case of L1 = 70 mm when the tapered rod was compared with the stepped tapered rod. Fig. 7 represents the shear stress distribution for each type of component having the overall length as 110 mm. A decreasing trend is observed for both the arrangement of rod with almost identical slopes. With the introduction of sudden step in the uniform cross-section, the shear stress induced was found to elevate. When tapering was introduced in the rod along with the sudden step, the shear stress was found to decrease. The minimum amount of shear stress was induced for the tapered rods without any stepping or abrupt change in the cross-section. Fig. 8 represents the variation of total deformation (on ordinate axis) for the various types of component (on abscissa) when the full length of the torsional rod was selected to be 110 mm. For the arrangement in which L1 was 60 mm and L2 was 50 mm, the total deformation was found to be maximum in the case of uniform stepped rod which decreased to the minimum value for the stepped tapered rod. The deformation for the tapered rod was observed to be moderate and had the value between that of uniform stepped rod and stepped tapered rod. On the contrary, in the case of the arrangement when L1 was selected as 70 mm and L2 was selected as 40 mm, a decreasing trend was exhibited by the rods with varying\nslope as the graph was steeper between uniform stepped rod and stepped tapered rod, which comparatively became less steep between the stepped taper and tapered torsional rod. Fig. 9 shows the strain energy of the cross-section when the full length was selected as 110 mm. For both the arrangements of torsional rod, it was observed that the slope was almost similar between the uniform stepped rod and stepped tapered rod. On the other hand, for the configuration when L1 was selected as 70 mm and L2 was selected as 40mm, the slope increased between stepped tapered rod and tapered rod, whereas when the L1 was selected as 60 mm and L2 was selected as 50 mm, the slope between the stepped tapered rod and tapered rod decreased.\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 20,2020 at 23:37:24 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.18-1.png", + "caption": "Figure 9.18 Elbow-down and elbow-up poses of an elbow manipulator.", + "texts": [ + "205) implies that if \ud835\udf0e1 = + 1 leads to \ud835\udf031, then \ud835\udf0e1 = \u2212 1 leads to \ud835\udf03\u20321 = \ud835\udf031 + \ud835\udf0b, which means that u\u20d7(1) 1 , i.e. the look-ahead direction of the manipulator, is reversed. (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e3 that arises in the process of finding \ud835\udf033 by using Eqs. (9.218)\u2013(9.220). The manipulator attains the same location of the pre-wrist point by assuming one of the two alternative poses that correspond to \ud835\udf0e3 = + 1 and \ud835\udf0e3 = \u2212 1. These poses are illustrated in Figure 9.18. The poses corresponding to \ud835\udf0e3 = + 1 and \ud835\udf0e3 = \u2212 1 are designated, respectively, as elbow-down pose and elbow-up pose. These designations are justified by Eq. (9.220), because it indicates that, for the same value of \ud835\udf093 = c\ud835\udf033, \ud835\udf033 becomes positive if \ud835\udf0e3 = + 1 and \ud835\udf033 becomes negative if \ud835\udf0e3 = \u2212 1. The corresponding values of \ud835\udf032 are determined by Eqs. (9.225) and (9.226). (c) Third Kind of Multiplicity The third kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the angles \ud835\udf035, \ud835\udf036, and \ud835\udf03234 by using Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure9-1.png", + "caption": "Fig. 9. Multibody model of the Waveboard: numbering of the bodies, generalized coordinates and body frames.", + "texts": [ + " In this section, the multibody model constructed by Garc\u00eda-Ag\u00fandez Blanco et al. [23] is used, considering the wheels as two rings in contact with the ground. The equations of motion are derived and a kinematic and inverse dynamics simulation is carried out in order to understand the maneuvering of the system. The multibody model consists of seven rigid bodies, corresponding with the rear and forward decks (bodies 2 and 5), back and front casters (bodies 3 and 6) and rear and forward wheels (bodies 4 and 7). Body 1 corresponds to the inertial frame. Fig. 9 shows the numbering of the bodies, generalized coordinates of the system and body frames. Coordinates x 2 , y 2 and z 2 locate the center of mass of body 2, and angles \u03b12 , \u03b22 and \u03b3 2 orientate the board in space. In order to take into account the twist of the forward deck with respect to the rear one, angle \u03b852 is used. Moreover, relative motion of the casters with respect to the decks is considered by means of the angles \u03b832 (for the rear caster) and \u03b865 (for the forward one). Lastly, angles \u03b8 and \u03b8 are used to describe the rotation of the wheels", + " Appendix A contains the 43 76 expressions of the origins of the body frames and the list of geometric and dynamic parameters of the model with their corresponding numerical values. On the other hand, the model presents eight constraints, four of them holonomic and another four non-holonomic. The holonomic constraints arise from forcing that the Waveboard must be in contact with the ground. In the first place, this requires that the Z -component of the position vectors of the contact points (denoted as R and F in Fig. 9 ) be zero: r R Z = 0 , r F Z = 0 . (23) Secondly, the contact point must correspond to the lower point of the wheel. This can be done by imposing that the tangent vectors to the contact points (denoted by t R and t F ) and the normal to the ground n are orthogonal: n \u00b7 t R = 0 , n \u00b7 t F = 0 . (24) In order to obtain the position vectors of the contact points r R and r F , and the corresponding tangent vectors, nongeneralized coordinates \u03be R and \u03be F are defined. These coordinates represent a counterclockwise rotation about the Y 4 and Y 7 axes, orientating axes X \u03beR and X \u03beF so that they point to the contact points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003272_j.promfg.2020.05.027-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003272_j.promfg.2020.05.027-Figure4-1.png", + "caption": "Figure 4. Configuration of the new L&T ultrasonic vibrator system.", + "texts": [ + " The defined wave path is a circular helix trajectory, which can be represented as follows: { \ud835\udc65\ud835\udc65(\ud835\udf03\ud835\udf03) = \ud835\udc37\ud835\udc37 2 \ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udc50\ud835\udf03\ud835\udf03 \ud835\udc66\ud835\udc66(\ud835\udf03\ud835\udf03) = \ud835\udc37\ud835\udc37 2 \ud835\udc50\ud835\udc50\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udc60\ud835\udf03\ud835\udf03 \ud835\udc67\ud835\udc67 = \ud835\udc3f\ud835\udc3f 2\ud835\udf0b\ud835\udf0b \ud835\udf03\ud835\udf03 , \ud835\udf03\ud835\udf03\ud835\udf03\ud835\udf03[0, \ud835\udf0b\ud835\udf0b] (1) Where D is the diameter of the circular helix, and the L is the length of the rods. Based on the wave path, six patterned helical rods were created as the final waveguide structure (see Figure 2). The diameter of the rods was designed as \u03a64.5 mm, which is determined by the maximum sustainable stress based on our FE analysis, as discussed in the next section. The design and configuration of the assembled new L&T vibrator are shown in Figure 4. To obtain high electromechanical conversion efficiency and reasonably stable performance, a bolt-clamped type Langevin transducer is used to generate the longitudinal vibration. The Lead Zirconate Titanate (PZT-8) piezoelectric ceramic rings are polarized along the thickness direction. The PZT-8 used in the design are commonly used in power ultrasonic applications. As shown in Figure 4, every two adjacent piezo-ceramics rings are oriented in the opposing direction, and among them, the copper electrodes are alternatively sandwiched. For safety concerns, the copper electrodes attached to the mass bolt are connected to the ground as negative terminals, while another pair of the copper electrodes are connected to the ultrasonic driver (see Figure 4). When sinusoidal voltages are applied to the PZT rings, all the PZT ceramic rings will synchronously generate longitudinal vibrations. To analyze and optimize the design parameters of the waveguide-based L&T vibrator, a finite element model is constructed for analyzing the resonance frequencies, mode shapes, and steady-state behaviors of the geometric design. The modeling software used was ANSYS 19.1 Workbench at our lab. The material used for the vibrator is aluminum alloy AlSi10Mg. The connecting bolt and back mass is made of 1045 steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003932_s40962-020-00549-5-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003932_s40962-020-00549-5-Figure6-1.png", + "caption": "Figure 6. Plaster molding setup.", + "texts": [ + " The model which predicts minimum temperature at the output is as follows: y \u00bc 1:4965x1 2:4922x2 \u00fe 0:027447x2 1 0:0004113x1x2 0:0016444x1x3 0:00048333x1x4 \u00fe 0:059556x1x5 \u00fe 0:0037906x2 2 0:019381x2x3 \u00fe 0:01541x2x4 \u00fe 0:25794x2x5 \u00fe 0:052841x2 3 \u00fe 0:0097348x3x4 0:0029764x2 4 0:59321x4x5 Eqn: 2 where y denotes the minimum temperature at the output, C; x1 up-profile radius 1, mm; x2 up-profile radius 2, mm; x3 down-profile radius 1, mm; x4 down-profile radius 2, mm; x5 thickness, mm. International Journal of Metalcasting The absolute error average value of this model is 0.12765 C, and its relative error average value is - 0.067878%. Simulation annealing is applied to this model for optimization. The execution result shows the minimum temperature at the output is - 208.75 C when A is 22.5, B is 122, C is 16, D is 79, and E is 4. The setup for manufacturing the plaster core is shown in Figure 6. As shown in Figure 6, the plaster molding setup is designed as a separation type so that only this part can be replaced when the type of wheel is updated. By the way, it is very difficult to make the top die1, top die2 and lower die by traditional methods. So in this work, these parts are manufactured using FDM equipment. The material of the part is PLA resin. The thermodynamic properties of PLA resin are shown in Table 3. International Journal of Metalcasting Figure 7a shows the FDM equipment (Ultimaker 2) for mold manufacturing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002051_j.procir.2016.02.003-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002051_j.procir.2016.02.003-Figure7-1.png", + "caption": "Fig. 7: Radial displacement of the inner ring for bearing components: a) considered ideally round, b) with a roundness deviation of 1 \u00b5m, c) with a roundness deviation of 3 \u00b5m, d) with roundness deviation of 5 \u00b5m. e) Comparison of the mean bearing clearance of each of these cases", + "texts": [ + " 5: Detection of the relevant vertices/edges on the outer ring for a certain roller point of contact area of interpentration a) b) Fig. 6: Raytrace algorithm a) without and b) with consideration of the opposing surface. This example focuses on the effect of the geometric deviation. Therefore different cases are investigated. First, only dimensional deviation is considered, i.e. every component is regarded to be ideally round. Then the maximal geometric deviation of every component is iteratively increased to a geometric deviation level of 1, 3 and 5 \u00b5m. The results are presented in the radar charts in Fig. 7. Each point represents a radial displacement of the inner ring in this direction. Hence, the distance between two opposing points represent the bearing clearance for this axis of displacement (consisting of one direction and its opposing direction). For each level 14 bearing clearances are determined, i.e. 28 directions are evaluated. The closer the graph of the radar charts is to a circle, the smaller is the variance of bearing clearance. Hence, the variance seems to increase with the geometric deviation. In contrary the bearing clearance seems to be quite constant for the first three cases (no geometric deviation, 1 \u00b5m and 3 \u00b5m geometric deviations) and suddenly drops for the last case (5 \u00b5m geometric deviations). This effect could also be seen in the lower part of Fig. 7. This diagram shows the mean bearing clearance for the different level of geometric deviation. The decrease of the mean bearing clearance due to an increased geometric deviation seems logic, as not only the amplitude of the waves but also the noise increases. This means that the surfaces of the bearing components become rougher wherefore the likelihood of two roughness peaks contacting each other is increasing. However, despite the geometric deviation the mean bearing clearance for a geometric deviation of 1 \u00b5m and 3 \u00b5m is slightly higher than for ideally round bearing components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001946_ecce.2014.6953971-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001946_ecce.2014.6953971-Figure5-1.png", + "caption": "Fig. 5 Discretization and calculation of the flux density in the stator.", + "texts": [ + "s ,,, qdyokeqdteethqdstator qdyokeqdteethqdsattor qdyokeqdteethqdstator iiciiciic iibiibiib iiaiiaiia (6) The parameters of material come from Flux 2d libraries and are given in the following table: Finally the iron losses model requires the calculation of the mapping of the coefficients astator, bstator and cstator for different pairs of currents (id,iq (5) ). By the interpolation of these mappings and the application of equation , it is possible to estimate iron losses for any pair of currents and speed, facilitating the coupling with others physical models. A complete description of this model is given in [13]. The magnetic nodal network discretizes the stator as described in the Fig. 5. The red points on Fig. 5 are the flux densities calculated in the middle of each zone discretized by reluctance on one pole. As the spatial evolution is assumed identical to the time evolution, therefore we have the following relationship: p Z T step Z step s s temporal s spatial \u22c5=\u21d2 \u22c5 = \u03c02 (7) With: Ts Z the sample time. s the slots number. With symmetry, we can determine the waveform on one period. The time evolutions of the flux density in the teeth and the yoke are described in Fig. 6. As the FDWs contain only 12 values by electrical period, an interpolation by cubic method with Matlab function interp1 has been applied as described in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003911_s42417-020-00269-4-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003911_s42417-020-00269-4-Figure1-1.png", + "caption": "Fig. 1 Pair of helical gears from an automotive gearbox", + "texts": [ + ":(0123456789) Keywords Rattle noise\u00a0\u00b7 Clutch\u00a0\u00b7 Gearbox\u00a0\u00b7 Conical springs\u00a0\u00b7 Vibration\u00a0\u00b7 Noise control The combination of an internal combustion engine (ICE) with a manual or semi-automatic gearbox, using a clutch system, is considered one of the most common configurations of vehicle powertrains. The engine crankshaft receives torque peaks, because of the combustion inside the chambers, which propagate throughout the rest of the powertrain. An issue that might appear in the transmission system, considering the combustion engine working principle, is rattle noise. This is a high-frequency sound caused by impact between the gearbox teeth due to backlash, as shown in the Fig.\u00a01. The torque peaks decelerate the conducting gears and the pairs loose contact with each other and when contact is reestablished the new acceleration causes the impact [1], through the use of an experimental test rig. The authors characterized the sensitivity of the noise threshold to the presence of gear backlash. The other authors used a test bench of a single-stage gearbox to describe the rattle noise level, varying the backlash and gear helix angle [2]. According to them, a reduction in the noise is associated with avoiding meshing impact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002262_978-3-319-27333-4-Figure2.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002262_978-3-319-27333-4-Figure2.1-1.png", + "caption": "Fig. 2.1 Schematic of circular journal bearing", + "texts": [ + "1007/978-3-319-27333-4_2 Chapter 2 Classifi cation of Non-circular Journal Bearings The chapter presents the reasons towards the development of non-circular journalbearing profi les and classifi cation of such profi les. The chapter also presents the basic mechanism of operation on which non-circular journal bearing works. In the last section of the chapter, the different regime of lubrication which may occur in various types of bearing has also been presented. The basic confi guration of the circular journal bearing consists of a journal that rotates relative to the bearing which is also known as bush (Fig. 2.1 ). Effi cient operation of such bearing requires the presence of a lubricant in the clearance space between the journal and the bush. In hydrodynamic lubrication, it is assumed that the fl uid does not slip at the interface with the bearing and journal surface. There exists a velocity gradient over the thickness of the fl uid, which depends upon the relative movement of bearing surfaces. There will be no pressure generation if the bearing surfaces are parallel or concentric, which means bearing could support any bearing load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001682_amm.789-790.226-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001682_amm.789-790.226-Figure1-1.png", + "caption": "Figure 1: Model completed in Adams", + "texts": [ + " This can increase the calculation efficiency and don\u2019t have ill effect on the precision of final results. In order to make sure the model does not have interference, the model should self-inspect in Pro/e. Saved the model as PARASOLID (*.x_t) format and imported it into Adams. This format doesn't have damage to the model and the accuracy can also meet the requirements. In Adams, Define the material properties of each unit. Then define the constraints, loads and motion correctly based on the RV reducer. The finished model was shown in figure 1. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (#68868371, Purdue University Libraries, West Lafayette, USA-30/07/16,16:14:52) Defining as the following: The influencing factors of transmission in the model. The influence to the transmission inaccuracy is composed of three aspects as deformation, backlash and the inaccuracy of manufacturing and assembly. The inaccuracy of manufacturing and assemble was affected by many factors and has complex combination which should conduct analyze separately" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.19-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.19-1.png", + "caption": "Figure 10.19 The (a) inward knee and (b) outward knee posture modes of the leg Lk.", + "texts": [ + "484) give \ud835\udf03\u2217k as follows without a new sign variable but on condition that c\ud835\udf19\u2032 k \u2260 0. \ud835\udf03\u2217k = \u2212atan2[\ud835\udf0e\u2032k(rk3 + b1s\ud835\udf03k), \ud835\udf0e\u2032k(rk1 \u2212 b1c\ud835\udf03k)] (10.490) Finally, \ud835\udf03\u2032k is obtained from Eq. (10.474) as \ud835\udf03\u2032k = \ud835\udf03\u2217k \u2212 \ud835\udf03k (10.491) (b) Multiplicity Analysis of Inverse Kinematics The inverse kinematic solution presented in Part (a) involves two sign variables \ud835\udf0ek and \ud835\udf0e\u2032k . Their effects are discussed below. The sign variable \ud835\udf0ek in Eq. (10.487) leads to two visually distinct posture modes for the leg Lk . These posture modes can be visualized by referring to Figure 10.19 and noting that \ud835\udf03k is the angle of \u2212\u2212\u2192BkCk and \ud835\udf03\u2218k is the angle of the projection of r\u20d7k = \u2212\u2212\u2192BkAk on the 1\u20133 plane of the frame 0k(Bk). Depending on the two values of \ud835\udf0ek , let \ud835\udf03+k = \ud835\udf03\u2218k + \ud835\udefek and \ud835\udf03\u2212k = \ud835\udf03\u2218k \u2212 \ud835\udefek . Then, the angles \ud835\udf03+k and \ud835\udf03\u2212k are ordered so that \ud835\udf03\u2212k < \ud835\udf03 \u2218 k < \ud835\udf03 + k (10.492) The posture modes with \ud835\udf03+k and \ud835\udf03\u2212k may, respectively, be designated as inward knee posture mode and outward knee posture mode. In the usual operations of the delta robot, though, the outward knee posture mode is preferred almost unexceptionally", + " For this manipulator, the first leg with \ud835\udefd1 = 0 is preferable for the sake of simplicity. Thus, the location of the tip point is determined as follows by using Eq. (10.473) p = u1(b0 \u2212 b7 + b1c\ud835\udf031) + u3(h0 \u2212 h7 \u2212 b1s\ud835\udf031) + b2eu\u03032\ud835\udf03 \u2217 1 eu\u03033\ud835\udf19 \u2032 1 u1 (10.531) (b) Multiplicity Analysis of Forward Kinematics The forward kinematic solution presented in Part (a) involves three sign variables \ud835\udf0e1, \ud835\udf0e2, and \ud835\udf0e3. They directly affect the angles \ud835\udf03\u22171 , \ud835\udf03\u22172 , and \ud835\udf03\u22173 , respectively. On the other hand, as implied by Figure 10.19, the sines of these angles (i.e. s\ud835\udf03\u22171 , s\ud835\udf03\u22172 , and s\ud835\udf03\u22173 ) indicate whether the moving platform is located below the knee points (i.e. C1, C2, and C3) or not. In other words, the moving platform will be below the knee point Ck if s\ud835\udf03\u2217k > 0. According to Eq. (10.528), such a pose can occur for the legs L2 and L3, i.e. for k = 2 and 3, if \ud835\udf0ek = sgn(nk3) (10.532) As for the first leg, Eq. (10.521) implies that such a pose can occur if \ud835\udf0e1 = sgn(K1) (10.533) However, as verified in Section 10.8.7, the moving platform remains always parallel to the fixed platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002003_cjme.2015.0713.092-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002003_cjme.2015.0713.092-Figure6-1.png", + "caption": "Fig. 6. Shaft system used in the experiments", + "texts": [ + " In this study, the whirling coefficient ranged from 2.370 to 5.418 as shown in Fig. 5(b). According to the theory by YANG, et al[21], Eq. (17) can be simplified as follows to explain the oil whip mechanism more clearly: 1 3 3 2 2 2 + 0.5 + , f f f f f f \u00bb \u00bb (18) where 2.3700.5\u2126 when \u03bd>0. The test rig includes a vertical shaft system with a 2 m long shaft supported by a roller bearing and a sliding bearing at the two ends of the shaft as shown in Fig. 6. The sliding bearing experienced the water whip. During operation, water was pumped from the inlet pipe at the bottom of the system, up through the bearing and then out the outlet pipe at the top. A grand packing seal was used to prevent water leakage at the top. The shaft working speed ranges from 0 r/min to 5000 r/min. Speed up and down (05000 r/min) tests were conducted with different flow rates to observe the water-film whip. The water flow rate was controlled by a valve to three flow velocities of 0 m/s, 2 m/s, 4 m/s. Four displacement sensors measured the shaft displacements at the bearing and at the middle of the shaft in both the X and Y directions as shown in Fig. 6 and 7. Fig. 8 shows the orbits of the axis center at typical rotational speeds with no water flow. The displacement amplitude of the shaft center is larger than that of the bearing center because the shaft center is farther from the bearing and has less displacement constraints. The two key rotational speeds are 2400 r/min and 4700 r/min. The orbits of both the shaft center and the bearing center increased suddenly at 2400 r/min, while only the shaft center orbit increases significantly at 4700 r/min" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002262_978-3-319-27333-4-Figure2.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002262_978-3-319-27333-4-Figure2.6-1.png", + "caption": "Fig. 2.6 Truly elliptical journal bearing", + "texts": [ + " The bearing so produced has a large clearance in the horizontal or split direction and a smaller clearance in the vertical direction. Elliptical journal bearings are slightly more stable toward the oil whip than the cylindrical bearings. In addition to this, elliptical journal bearing runs cooler than a cylindrical bearing because of the larger horizontal clearance for the same vertical clearance. Some other non-circular journal bearing confi gurations; Three-lobe journal bearing (symmetrical and asymmetrical, Fig. 2.5 ), Elliptical journal bearing (profi le is elliptic in cross-section, Fig. 2.6 ), and orthogonally displaced journal bearing (vertical offset, Fig. 2.7 ); have also been shown below. 2 Classifi cation of Non-circular Journal Bearings 7 e RL RL RL Fig. 2.5 Symmetrical three-lobe bearing In earlier works, the bearing performance parameters have been computed by solving the Reynolds equation only. Over the years, many researchers have proposed number of mathematical models. A more realistic thermohydrodynamic 2.3 Methods of Analysis 9 ( THD ) model for bearing analysis has been developed which treats the viscosity as a function of both the temperature and pressure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003133_icoa49421.2020.9094495-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003133_icoa49421.2020.9094495-Figure1-1.png", + "caption": "Fig 1: Motorcycle\u2019s tire mesh model", + "texts": [], + "surrounding_texts": [ + "Keywords\u2014 Bond graph, Tire, 20-Sim simulation, Strictness road, Piezoelectric model, Harvester System.\nI. INTRODUCTION\nActually the energy transitions enforce researchers to develop new, strong and reliable models for producing and storing electrical energy from renewable sources, outside the concerns of standard renewable energy sources, from which it is necessary to seek to develop or adapt existing systems to these criteria. In an original process, this paper will examine the possibility of the recovery of the vibratory energy dissipated by the motorcycle tires. To decrease motorcycle vibration by dissipating this energy is the purpose of the shock absorber in a motorcycle's suspension system, the dissipated energy can be employed for alimented motorcycle's accessories with the piezoelectric beam pasted on the motorcycle tire. Based on the fact that the motorcycle's suspension system is a dynamic device invent to sleek out or dispirit the shock impulse, and wear kinetic energy. In-vehicle design, it reduces the effect of moving over the rasping ground, leading to developed ride quality, and increase in driving comfort due to substantially disturbances reduced. The tunability of the vehicle travel on a level road with the wheels strike a bump, the suspension's spring is compressed quickly. This compressed spring will be modeled by a transversal force applied to attempt to return to its normal loaded height and harvest the vibration energy to electric power, in so doing, will rebound past its normal length. The weight of the motorcycle causes the corps to be lifted and harvest a large will\nthen push the suspension's spring down below its normal loaded height. This method ensures to the suspension-tire wheels, a safety motion and a little less each time when the tire is in Virage, this process is uncontrolled for suspension and tire at the same time until the up-and-down movement finally stops. The Bond Graph model is allowed to it will not only cause a max electric energy ride and will make handling of the motorcycle very comfortable. An important factor in modern electrical motorcycle fuel consumption, tire resistance is considered a major component of 2-wheels energy loss [1]. The tire deflection and deformation while rolling on the road caused friction on the road with piezoelectric material hysteresis and tire rubber, the tire structure is the primary contributor to energy dissipation. When the tire is moved on a flat surface, energy losses occur due to the contact patch being deformed in the limited condition according to the property and flexibility of the tire's structure internal, rubber, sidewall, and tread [2]. On the other hand, the tread element is deflected in the horizontal direction for the tire rolling on an uneven road, and rolling resistance will be generated by the longitudinal force versus motion because of the road variability. Recently, most of the researches have concentrated on the rolling strength of the tire at steady-state conditions [3], but in this document, special regard is paid to the energy loss generated in the tire due to the longitudinal force resulting from the tire rolling on an uneven road in order to harvest this lost energy and converted to electric useful power, this dissipated energy through processes such as plasticity [4], viscoelasticity or creep, the energy dissipation because of tire on the road sliding friction this is mean that the energy dissipation generated by the longitudinal force because of the road variability is principal component of energy loss [5], which could be harvested as electric energy from external road work [6-7], and is the main source for autonomous motorcycle electrical consumption. This work was divided into five sections: in section 2, a description of the motorcycle dynamic connected to the piezoelectric harvester device, and tire characteristics are calculated. In section 3, the motorcycle tire-road interaction proposed model is shown with the proposed mechatronic modeling. Simulation results and discussion are exposed in section 4. the piezoelectric beam bond graph model for the motorcycle tire rolling system is presented in cornering motion. Section 5 highlights the conclusions and perspectives.\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 18,2020 at 05:31:31 UTC from IEEE Xplore. Restrictions apply.", + "II. MECHATRONIC MODELING APPROACH\nCurrently, a lot of different active embedded devices added on the modern motorcycle needed new harvester energy systems which increase electric energy for motorcycle accessories [6-7], recently they become a standard on all motorcycles in order to increases the lifespan of the batteries.\nThis mechatronic approach allows the development and analysis of motorcycle electric power needed for embedded device, the availability of a suitable [8] motorcycle power is important to optimize this need, so we propose the use of a piezoelectric patch on the motorcycle tire without changing the tire's characteristics in order to test and design the estimation model characteristic before moving to complete system tests.\nIn this case, the proposed mechatronic studies, modeling and control for the motorcycle tire dynamic variable allows an optimal concept for this system. In order to ensure a good modeling test we are based on the 20-Sim software which give us the tire dynamic characteristics Fig.2. The motorcycle tire corps is modeled by spring-damper-mass systems, with real characteristics in the improved Bond Graph model Figs.3.4. The improved model input variables are the road profile and the motorcycle steer front wheel. The road profile applied forces on\nthe motorcycle tire are expanded according to a real road profile. In order to follow a real trajectory we impose in the BG model a virtual driver's masse ,the road longitudinal and lateral forces values, the modeling result show a perfect follow-up.= (1)= (2)= + \u0307\nCurrently, the global energy crisis imposes an additional burden on industry developers. However, we improve the piezoelectric harvester to investigate the vibration energy, the piezoelectric system characteristics noise effects are modeled by a hysteresis behavior, in terms of accessories switching frequency, to find the resonate frequency for the piezoelectric beam, this frequency is given on the model performance and it is necessary to be reachable on all systems. Hence, we chose to analyze this new piezoelectric harvester system which does directly rely on a motorcycle tire body by mechatronic modeling which is used to identified the input/output parameters value of the tire dynamic and the piezoelectric system in order to not affect the tire characteristic for harvest the road vibration energy, so to investigate its suitability to be employed in model-based concepts. In order to reduce system complexity and concepting time, we develop the tire quality and we miniaturize the piezoelectric batch design ,the simulation result made in the 20- Sim environment expose the system performance. This section details the piezoelectric harvester system analytic model and the tire dynamic behavior equs1.2.3. for doing that we are based on different reviews for harvester system schemes that are used for vehicle suspension systems on electric vehicles [7] in order to improve it and use it on the motorcycle tire. As the present road conditions cannot be assumed to be famous, in modeling design we adopted a novel Bond Graph tools this choice, for this case, give the set-point computation which is calculated assuming a road-friction coefficient value \u03bc = 0.3. This value, for tire components applications, can be regarded as a good value, as it is effectively a low friction condition but, if reached in reality, would not compromise rideability. Piezoelectric harvester systems that are modeled and simulated are shown in Fig5.\nA bond graph model is improved for a tire-suspension motorcycle system considering lateral system resistance and longitudinal road forces. Furthermore the 20-Sim simulation results explain the system performance and quality, we compared those result to the motorcycle estimation energy consumption we deduce that the harvested energy is higher than the motorcycle needed energy, so we can provide an additional simulation results obtained from MATLAB simulation of the solid model of the motorcycle tire prototype. In this case, our approach is to optimize the time and energy lost during the analyze. We know that complex systems require a lot of skills. Such mechatronic modeling is interesting to allow the seat point of displacement or acceleration applied (as equivalent force) and assuming that there is no disturbance. Moreover, theoretically, to calculate the equations of states of a dynamic system [9], it takes a huge amount of work.\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 18,2020 at 05:31:31 UTC from IEEE Xplore. Restrictions apply.", + "This modeling has made it possible to respond to needs, it will give us the force vectors and the flow vectors using a symbolic form of the integrated causal link which show in the (Fig. 7) [8]. We apply our approach for n modes, the relationship between the electrical behavior and the motion equation is given using a bond graph modeling [9].\nIII. PIEZOELECTRIC BEAMS ON MOTORCYCLE TIRE\nGenerally, the piezoelectric parameters are issue from on the IEEE data value, those beams characteristics expose the nature of the harvester system developed, the power harvested requirements will be available on tire device or motorcycle tire wear. So, this system is appropriate for a smaller device which connected to the road or unknown terrain.\nThis piezoelectric beam model which is connected to the motorcycle tire allows produce the electric power by the motorcycle motion in the irregular road by convert the vibration caused by the wheel motion to electrical power those variables were taken into account and discussed in the bond graph proposed design.\nFor this mechatronic modeling we consider the motorcycle tire steering using a prototype example for independent or dependent inputs, the different matrix values are calculated by the identification strategy in MATLAB, the simulation Figs. 5,6 ,show that the system is stable in the present configuration. A 20-Sim simulation is followed by a full motorcycle complete model. A tire Bond Graph model is sufficient to examine and harvest the body mode of vibrations; however, if one intends to develop the vibrating model of a system to include modes of system vibrations. It can also provide a piece of comprehensive knowledge about the road roughness inputs for the Bond Graph model to the suspension and responses of these systems.\nThe electrical power needs in order to concept an autonomous motorcycle is calculated by this bond graph model Fig. 5, consider the fundamental principia of a standard piezoelectric harvester system modeled in [6], this section aims to analyze the motorcycle behavior, the proposed energy harvesting system has very rich nonlinear dynamics [10], including chaos, perioddoubling motion, and others [11-12],this is way we use the bond graph tools to take into accent the different perturbation. = ( . ) ; = (4 Moreover, this system can harvest a large amount of energy which is a chaotic attractor, it can be clearly seen if they work dimensions, structure, so the electrical output of a piezoelectric harvester is important and input energy remains unchanged, the Eq (4) is dominated by the (d\u00b7g) term. Furthermore, by varying a single parameter, a chaotic attractor is detected in the novel system\u2019s overcome this problem we proposed this Bond Graph model to extract the system variable values easily.\nIV. SIMULATION AND DISCUSSION\nThe simulation result for the proposed bond graph model expose the travelling velocity which is considered as a constant value 70km/h, in order to have a clear view of our studying system we have present this calculation technique of singular perturbation, it is when the dynamics of the system is well separated.\n( ) = \u2217 \u222b \u2016 0 \u2212 \u2016\u00b2 (5 ( ) = \u2217 \u222b 1\u0307 \u2212 \u0307 \u00b2 (6\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 18,2020 at 05:31:31 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_34_0001647_0954406215625677-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001647_0954406215625677-Figure3-1.png", + "caption": "Figure 3. Schematic of the computational domain of numerical analysis about finger seal leakage.", + "texts": [ + " In other words, the leakage gap hi(t) between the finger sticks in the ith finger element and rotor at time t could be defined as hi\u00f0t\u00de \u00bc xi\u00f0t\u00de yi\u00f0t\u00de \u00f014\u00de where xi(t) and yi(t) are finger stick\u2019s displacement response and the corresponding displacement excitation of rotor at time t. The former is achieved by solving equations (1) and (2), and the latter is given by equation (12). To consider the effect of the compressibility of the fluid, the leakage rate of the finger seal is calculated by means of the numerical analysis method of computational fluid mechanics. The numerical analysis computational domain about the flow field of the finger seal leakage is built, and its schematic is shown in Figure 3. The x, y, and z directions are the circumferential, radial, and at The University of Melbourne Libraries on June 4, 2016pic.sagepub.comDownloaded from axial directions of the finger seal, respectively. The leakage gaps between each finger stick in some finger elements and the rotor at one time are described in Figure 3, and the fluid leakage is formed in the gap. It should be noted that the shape and size of the flow field computational domain are different at different times as well. Therefore, the calculation of the finger seal leakage needs to superimpose the leakages corresponding to each time in a motion period. That is to say, it is important to calculate the fluid leakage in the computational domain corresponding to each time. If the time steps of a motion period of the rotor are set as f, then the computational domain, which is shown in Figure 3, is given, the size in x direction is pDr/f; the size in y direction is the leakage gaps of the finger sticks in each finger element and the rotor. It should be noted that the leakage rates are different under different times and finger elements and are confirmed based on equation (14); the size in z direction is the thickness of the finger seal, which is composed by n-layer finger elements. Computational fluid dynamics (CFD) numerical calculation is carried out based on the computational domain which is shown in Figure 3, and its boundary conditions are defined as follows: the high pressure PH and low pressure PL are applied to boundaries 1 and 2, respectively; boundary 3 is the finger foot surface, therefore the static no-slip wall condition is applied to this boundary; boundary 4 is the circumferential surface of the rotor and the kinetic no-slip wall is applied to this boundary whose velocity is the linear velocity of rotor surface; cyclic symmetry boundary conditions are applied to boundaries 5 and 6. The flow field computational domain, which is shown in Figure 3, is dispersed using the structured hexahedral grids, the numbers of the grid in the x and y directions are M and K, respectively. The fluid is treated as compressible fluid in the numerical calculation, and the governing equations of air flow field are dispersed by the high order upwind format, then the velocity field of the fluid in the flow field is calculated using SIMPLE algorithm. The outlet of the flow field is chosen as the location of the leakage calculation. The velocity of arbitrary fluid microunit in the z direction is wmk(t), and the fluid density in the fluid microunit is mk(t), then the mass flow rates of all fluid microunits are superimposed, therefore the leakage rate of the finger seal at time t could be defined as q\u00f0t\u00de \u00bc XM m\u00bc1 XK k\u00bc1 wmk\u00f0t\u00de mk\u00f0t\u00de lx\u00f0t\u00de ly\u00f0t\u00de \u00f015\u00de where lx(t) and ly(t) are the sizes of the fluid microunits in the x and y directions at time t" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002272_2016-01-1958-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002272_2016-01-1958-Figure8-1.png", + "caption": "Figure 8. Distortion corresponding to concavity at P1, P2, P3, and P4", + "texts": [ + " Maximum distortion occurs at the shield perimeter due to the size being both thin and large. It is important within this figure to focus on the outer ring distortion as key areas, that is, mounting location and raceways. Four points, outboard seal fitting point, P1, outboard raceway ball contact point, P2, inboard raceway ball contact point, P3, and inboard seal fitting point, P4, were selected from the outer ring because raceway contact points and seal mounting points are important for bearing performance. Figure 8 shows the distortion of the aluminum knuckle in correspondence to the concavity at P1, P2, P3 and P4. Distortion shapes of P1, P2, P3 and P4 resemble a triangle, polygon, and inverted triangle. Distortion shapes of P1 and P2 were slightly similar, while P3 differed and P4 were inverse of P1 and P2. The difference can be attributed to the fact that P1 and P2 are located on the outboard side, while P3 and P4 are located on the inboard side. In addition, distortion shapes varied with concavity changes of 0 \u03bcm to 50 \u03bcm in 10 \u03bcm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003126_iraset48871.2020.9092240-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003126_iraset48871.2020.9092240-Figure2-1.png", + "caption": "Fig. 2. Mesh curve of taper rod.", + "texts": [], + "surrounding_texts": [ + "978-1-7281-4979-0/20/$31.00 \u00a92020 IEEE\nKeywords\u2014Torsion rod, Tapered rod, vibration, stepped tapered rod, ANSYS, CATIA, Rod, Automobile, Stepped Rod.\nI. INTRODUCTION Suspension systems are implemented on the basis of road profile or vehicle purpose. Based on light, medium and heavy-duty operations, various systems are employed. Coil springs are used for medium duty and flexible operations, whereas leaf springs are used for heavy load vehicles such as trucks and buses. Hatch back cars use either coil springs or torsional springs. Coil spring provide variable spring rate which means load required for per unit deformation increases as the springs get compressed. Also, coil spring increases the front part volume of the cars. On the other hand, torsional rod is attached out of the chassis where one end of the torsional rod is attached to the chassis and the other end is keyed to the stub axle of the wheel. Vibration, by its definition, is the displacement of the body from its equilibrium position. In the case of torsional rod, torque is experienced by it due to bumping of tyres on pits and twist, and torsion rod regain its equilibrium position when the tyre comes out of the pit. It is also important to mention that a preload is usually applied at the torsional rod which aids in maintaining the ride height of the vehicle.\nI. LITERATURE REVIEW For better comfort and good riding experience, researchers are continuously experimenting and analyzing various suspension systems. Radhakrishnan et al. [1] analyzed the behavior of stepped torsional rod for various parameters. Alizade et al. [2] analyzed the non-linear torsional behavior of pre-stretched rod and observed that non-linear theory coincides with linear theory of torsional vibrations for thin elastic and pre-stretched rod. Lu et al. [3] analyzed multistepped beams with composite element theory, which is the derivative of finite element method and analytical theory method used together. Concepts of Strength of Material and Torsion were also utilized in this work [4,5]. Radovanovic et al. [6] employed continuous mass transfer matrix method\nto determine the natural frequency of bodies which were attached to elastic beams. Matijevic\u00b4 et al. [7] successfully investigated the influence of automobile vibrations on human body by utilizing modern signal analyzing technique. Vijayan et al. [8,9] studied the methodologies which were effective for isolating the vibrations caused by a system using passive approach and topology optimization.\nII. EXPERIMENTAL METHODOLOGY In the present investigation, CATIA V5R20 was used for modeling of torsional rods and the models created on CATIA were later analyzed using ANSYS 2019R3. CATIA V5R20 stands for computer aided three-dimensional interactive application version 5 and revision 20. ANSYS stands for analysis systems. ANSYS utilizes the concept of finite element method, in which a component is meshed into finite number of elements and effect of load or boundary condition is observed on each element. Junction of every element is known as node. Due to the constraint of limited number of pages, only results are shown for each type of component. Table 1 illustrates the various parameters for modeling of torsional rod as proposed by Radhakrishnan et al. [1]. Using the aforementioned parameters, four models of torsional rod with different dimensions was prepared for each of the following: (i) stepped rod, (ii) stepped taper rod, and (iii) taper rod. For stepped rod, the geometry of the various models was as follows: In the case of first, second, third and fourth models, the end diameters were selected as 4.37 mm and 4.21 mm, 4.45 mm and 4.11 mm, 4.37 mm and 3.80 mm, and 4.21 mm and 4.02 mm respectively, which were separated by the rod length of 130 mm, 130 mm, 110 mm, and 110 mm respectively, and was stepped at a distance of 70 mm, 75 mm, 70 mm and 60 mm from the bigger diameter end respectively.\nFor stepped taper rod, the geometry of the various models was as follows: in the case of first, second, third and fourth model, the bigger end diameter was selected as 4.45 mm, 4.45 mm, 4.37 mm and 4.37 mm respectively which tapered upto an axial distance of 70 mm, 75 mm, 70 mm and 60 mm respectively, and resulted into a gradual reduction of diameter to 4.37 mm, 4.37 mm, 4.21 mm and 4.21 mm respectively, which was followed by a sudden reduction of diameter to 4.21 mm, 4.21 mm, 4.02 mm and 4.02 mm respectively at the same distance which was subsequently followed by tapering to the final smaller diameter of 4.11 mm, 4.11 mm, 4.02 mm, and 4.02 mm respectively, upto the axial distance of 130 mm, 130 mm, 110 mm, and 110 mm from the bigger diameter end respectively.\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 20,2020 at 23:37:24 UTC from IEEE Xplore. Restrictions apply.", + "For taper rod models, the geometry of the various models was as follows: in the case of first, second, third and fourth model, the end diameters were selected as 4.37 mm and 4.21 mm, 4.45 mm and 4.11 mm, 4.37 mm and 3.80 mm, and 4.21 mm and 4.02 mm respectively, which were separated by the rod length of 130 mm, 130 mm, 110 mm, and 110 mm respectively. It is important to note that the taper rod models had no sudden change in cross-sectional are resulting into gradual change in diameter from larger diameter to smaller diameter along the axial length, whereas in the case of stepped taper rod, a steep change in the cross-sectional area was observed at an axial distance of 70 mm, 75 mm, 70 mm and 60 mm from the bigger diameter end in the case of first, second, third and fourth model respectively.\nIII. CALCULATION The material properties and length used in this study are mentioned below which has been obtained from the investigation performed by Radhakrishnan et al. [1]:\nFull length of the tapered bar (L) = 130 mm, and 110 mm. Maximum Torque (M) = 17600 N mm. Modulus of Rigidity (G) = 80 X 103 N/ mm2. Young\u2019s Modulus (E) = 207 X 103 N/mm2. The equation of torsion is [4]:\nL The equation of torsion strain energy is [4]:\nIV. ANSYS ANALYSIS Models prepared using CATIA were imported in ANSYS software for analysis. Material properties as mentioned in calculation section were assigned to each model which were meshed (default) later. The mesh generated for the uniform stepped rod, taper rod, and stepped taper rod has been shown in fig. 1, 2 and 3, and their corresponding meshing details have been illustrated in table 2, 3 and 4 respectively. Fixed support was applied at the larger diameter end and torque was applied at smaller diameter end. In all the cases of torsional rod including uniform stepped rod, taper rod and stepped taper rod, the triangular surface meshing was opted to perform finite element analysis. In the case of uniform stepped rod, the total number of elements\nselected varied between 236 to 412 whereas the total number of nodes used varied between 598 to 904, both of which dependent on the rod geometry (table 2). In the case of taper rod, the total number of elements was selected as 210 for all the different rod dimensions whereas the total number of nodes were selected as 1108 (table 3). While analyzing the stepped taper rod, the total number of elements selected varied between 238 to 409 whereas the total number of nodes used varied between 623 to 915, both of which dependent on the rod geometry (table 4).\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 20,2020 at 23:37:24 UTC from IEEE Xplore. Restrictions apply.", + "V. RESULTS AND DISCUSSION The various results obtained by utilizing the CATIA and ANSYS for the different types of torsional rod have been plotted and summarized in the subsequent sections. Fig. 4 illustrates the behavior of shear stress (on ordinate axis) with the various torsional rod components (uniform stepped, taper stepped and tapered rod on abscissa) for the torsional loading. For arrangement with L1 = 75 mm and L2 = 55 mm, it is observed that the shear stress has a decreasing trend with a constant slope in the following order: stepped rod > stepped taper rod > taper rod. On the contrary, for the arrangement with L1 = 70 mm and L2 = 60 mm, the slope\nfirst slightly increases between stepped and stepped tapered which is followed by a decrease in the value between stepped tapered rod and tapered rod. No significant change in the slope is observed for rod having step in cross section. Contrary to this, in the case of taper rod with no step in their cross-section, shear stress is observed to be relatively lesser. Fig. 5 represents the behavior of cross-sectional deformation for the rod of 130 mm full length under the application of torque. For the arrangement of rod with L1 = 70 mm and L2 = 60 mm, the total deformation was observed to decrease while moving from uniform stepped rod to stepped tapered rod whereas it slightly increased while moving from stepped tapered rod to tapered rod. A continuously decreasing trend was observed for the rod having L1 = 75 mm and L2 = 55 mm with almost constant slope which was quite similar to the behavior of shear stress in the same rod. The sudden step in the uniform cross-section increases the deformation, whereas the step in the cross-section over the tapered rod was found to undergo comparatively lesser deformation. At the same time, it was also observed that the tapered rod and stepped tapered rod had almost identical amount of deformation. It may be concluded from the above observations that the step in the cross-section doesn\u2019t play a vital role in the case of tapered rod to provide resistance against deformation.\nAuthorized licensed use limited to: University of Exeter. Downloaded on June 20,2020 at 23:37:24 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_34_0001482_1.4029187-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001482_1.4029187-Figure5-1.png", + "caption": "Fig. 5 A dead center position for a symmetrical eight-bar linkage and its equivalent linkage", + "texts": [ + " The three passive joints I26, I65, and I15 of the type II equivalent four-bar linkage I12I26I65I15 are automatically collinear. The same result as Ref. [4] can be obtained with the proposed method. Another case is that if one of the type II equivalent four-bar linkage is at dead center positions, the whole linkage is at dead center positions while the type I four-bar linkage is not necessarily collinear, as shown in Fig. 16 in Ref. [4]. Example 6 will explain this observation. Example 4. Dead center positions for a symmetric eight-bar linkage. A symmetric eight-bar linkage of Fig. 5 consists of ternary links (or triads) A0AA0 and B0BB0, one rigid quaternary coupler (or tetrad) CDEF, and binary links AC, A0D, BE, and B0F. The equivalent four-bar linkage is A0GHB0 [13], which is obtained from the original four links 1, 2, 5, and 8 and formed by the four corresponding instant centers I12, I25, I58, and I81 (or I18). Point G (or I25) is the intersection point of AC and A0D, and H (or I58) the intersection point of BE and B0F. The coupler GH of the equivalent four-bar linkage may be regarded as being rigidly attached to CDEF, and A0G to A0AA0, and B0H to B0BB0. According to the criterion, in this case, the dead center position occurs when the three passive joints I23, I34, and I14 lie a common line, as shown in Fig. 5. Example 5. Dead center positions for the double butterfly eightbar linkage. Foster and Pennock [18] proposed a geometric method to locate the secondary instant centers of the double butterfly eight-bar linkage (Fig. 6), which consists of 3 five-bar loops. The double butterfly linkage in Fig. 6 has the similar dimensions as Table 1 in Ref. [18]. The dead center positions of the double butterfly linkage can be determined by any of its equivalent linkages. Let links 1, 2, 3, and 7 as the four original links from the double butterfly linkage and they are put in order as the reference link, the input link, the couple link, and the output link" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002627_icmic.2016.7804248-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002627_icmic.2016.7804248-Figure2-1.png", + "caption": "Fig. 2. 12-segments body model with 8-DOF considered in this work.", + "texts": [ + " Considering at first only flexion/extension movements, the rower body can be assimilated to an 11-degree-of-freedom DOF kinematic chain that can be reduced to 8-DOF; the feet being fixed to the stretchers and the rower sitting on a sliding seat, the lower limbs movements are identical, so only one side for lower limbs movement. As the two levers can be operated independently, the upper limb right and left side movements may be different. This allows one to definite eight basic rotations illustrated in Fig.2 and related to ankle (RA), knee (RK), hip (RH), pelvis (RP), left and right shoulders (RSL, RSR), and elbows (REL, RER). RSL and RSR rotation axis located at the upper limb extremities correspond to the joint linking the rower hands and rowing machine handles. They are not considered in this work. In a second time, one can notice that pushing/pulling the handles lead the levers to rotate around their axes, and the distance between the hands varies resulting in upper limbs adduction/abduction. Thus, rowing cannot be considered as a movement taking place only in sagittal plane; and one more degree-of-freedom should be considered at both shoulders and wrists joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003515_j.ijsolstr.2020.08.022-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003515_j.ijsolstr.2020.08.022-Figure3-1.png", + "caption": "Fig. 3. A diagram of the target configuration.", + "texts": [ + " For the initial configuration shown in Fig. 1, it can be determined that h=H \u00bc 0:29509 corresponding to h \u00bc p=8. The member axial stiffness mk of all the bars in the tensegrity is 1 107 N, and that is 1 105 N for all the cables. It can be determined that the tensegrity has a unique mode of self-stress state (Pellegrino and Calladine, 1986). The axial forces of some typical members at the initial configuration x0 are listed in Table 1. The tensegrity will be deployed to a target configuration xm (Fig. 3) in which some DOFs of joints A6, A8, B6, B8, C6, C8, D6 and D8, which are all on the top surfaces of the modules, are specified to reach the given positions. Comparing xm to x0, the xcoordinates of these 8 joints are at increments of +0.8 m, +0.8 m, +1.6 m, +1.6 m, +2.4 m, +2.4 m, +3.2 m and +3.2 m. In addition, joints D6 and D8 also have y-coordinate increments of +3.2 m. The locations of the other unconstrained DOFs, i.e., the unspecified DOFs, at xm are not limited. All 72 cables are adopted as active members, while the 16 bars are passive members with constant rest lengths" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure4.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure4.2-1.png", + "caption": "Figure 4.2 A manipulator with six revolute joints.", + "texts": [ + "78) As for the relative angular acceleration, it is found as follows by noting that Dan\u20d7 = 0\u20d7 because n\u20d7 is fixed with respect to a. ?\u20d7?b\u2215a = Da?\u20d7?b\u2215a = Da(?\u0307?n\u20d7) = ?\u0308?n\u20d7 + ?\u0307?(Dan\u20d7) \u21d2 ?\u20d7?b\u2215a = ?\u0308?n\u20d7 (4.79) Equations (4.78) and (4.79) happen to be very convenient in many practical situations, because the members of a mechanical system are connected mostly with single axis joints (i.e. revolute, prismatic, cylindrical, and screw joints) and the axis of such a joint appears fixed with respect to the members it connects. Consider the serial manipulator shown in Figure 4.2. It comprises six revolute joints. The unit vector along the axis of the kth joint is n\u20d7k . The reference frame attached to the link k is k . The base is the zeroth link (i.e. 0). The last link 6 is the end-effector, which is a gripper for this manipulator. The orientations of the successive links are described as shown below. 0 rot(n\u20d71,\ud835\udf031)\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u21921 rot(n\u20d72,\ud835\udf032)\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u21922 rot(n\u20d73,\ud835\udf033)\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u21923 \u00b7 \u00b7 \u00b75 rot(n\u20d76,\ud835\udf036)\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u21926 (4.80) In Description (4.80), the unit vector n\u20d7k of the joint k between the links k\u22121 and k appears fixed with respect to k\u22121 and k " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure2-1.png", + "caption": "FIG. 2. A schematic drawing of a diametrically creased disk (a) with a fold angle of \u03b2, disk radius R and hole radius r. A creased disk is shown mounted in the test fixture in (b). Stabilizing pins are located at the outer edges of the crease and fit into a slot in the test fixture. An additional stabilizing pin extends from the tip of the indenter and passes through the central hole. The disk is balanced on two cone pointed bolts positioned beneath the crease.", + "texts": [ + " EXPERIMENTS Tests were performed using 100 mm diameter disks cut from 0.10-mm-thick stainless steel sheet (AISI 301). To avoid the influence of a stress singularity at the vertex, all disks had a central hole with a diameter of 3 mm. Radial creases were formed by pressing the disks against a stiff rubber sheet using a die press. For simplicity, experiments were limited to disks with two or three equally spaced radial creases. Each disk was 3D scanned to obtain the initial fold angle across the crease, \u03b2 [see Fig. 2(a)]. In the case of a diametrical crease, a best-fit plane was found for the face on each side of the crease and the angle between their normals used as the fold angle. In the case of three radial creases, the scanned shape was cut by a plane perpendicular to each crease line at its midpoint. A best fit curve was obtained for each side of the crosssection and the angle at each intersection measured. The average of these angles is used as the crease angle for the disk. During testing the disks were supported along the crease line by cone-pointed bolts located 40 mm from the center. Since the diametrically creased disk is supported at only two points it is inherently unstable. Therefore, at each end of the crease, stabilizing pins were attached which fit into a slot in the test fixture to stabilise the disk during testing, as shown in Fig. 2(b). The disks were loaded using an indenter with a 4-mm-diameter point. From this point a 2-mm-diameter smooth rod is attached and passed through the central hole in the disk. This improves the stability of the test by preventing the disk from sliding off the supports. The three-crease disks were supported along the creases by three cone-pointed bolts, located 40 mm from the center. Stabilizing pins were not needed at the end of the creases since this geometry is stable. The three-crease disks were loaded with the same indenter and stabilizing pin as the diametrically creased disks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001676_amm.823.161-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001676_amm.823.161-Figure4-1.png", + "caption": "Fig. 4. a) The metal spherical prosthesis b) Cup polyethylene prosthesis c)The elastic ring prosthesis", + "texts": [], + "surrounding_texts": [ + "One of the most powerful tools usually used to analyze the stresses and displacements of human musculoskeletal system, but, also, of human joints-prosthesis assemblies is Finite Element Method [1-6]. Application of the finite element (FE) method for joint replacement was first completed using a femoral hip prosthesis in the late 1970s [1]. The first anatomically correct three-dimensional femur model with hip prosthesis was completed by Huiskes et al. [2]. The last decade has seen increased mesh resolution simulations, non-linear material models, and contact analysis to increase realism [3-11]. The Virtual Model of the Hip Prosthesis To determine the geometric parameters of the prosthesis components made by Groupe Lepine, we used direct measurement method. Also, we identified the spatial profiles of each component. Having determined the size of geometric elements of hip prosthesis, SolidWorks [7], parameterized software for three-dimensional models, was used. Using specific commands, sketch drawing plane is obtained (Fig. 2). Using the Insert command Boss / Base Extrude Depth parameter and assigning (from left window shape-defining) the amount of 9 mm solid is obtained. On the basis form we attached another additional form. In Fig. 3, a) the main steps of defining the stem model of hip prosthesis are shown. Finally, we have obtained the stem model of hip prosthesis shown in Fig. 3, b). All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 130.237.29.138, Kungliga Tekniska Hogskolan, Stockholm, Sweden-22/02/16,19:14:10) Assembling these elements, based on simple geometric constraints, it was obtained the final model of classic hip prosthesis (Fig. 5.). The Virtual Model of the Hip Joint Prosthesis To obtain the virtual model of the prosthetic hip joint, the surgical techniques were studied. We started from the virtual biomechanical system of both human legs elaborated in [13]. The implantation is preceded by removal of bone and cartilage in both affected bones: femur bone and pelvic bone. In a first phase, we removed the neck and the femoral head using a phisiological plane inclined at 45\u00b0 to the sagittal plane. The femur model prepared for implantation is similar to that shown in Fig. 6. To have a correctly entering of the stem prosthesis, the intramedullary canal was virtual widened, assuming a similar shape to that of the femoral stem (Fig. 6). In a sharp pelvic bone was performed similar spherical outer surface as the polyethylene cup (Fig. 7). Finally, using movement and simple geometric constraints, but based on orthopedic information (sagittal angle of 45\u00b0 and anteversion angle of 12\u00b0 for the polyethylene cup) was obtained prosthetic hip joint model, shown in Figure 7." + ] + }, + { + "image_filename": "designv11_34_0003203_j.est.2020.101417-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003203_j.est.2020.101417-Figure3-1.png", + "caption": "Fig. 3. The exploded view of the meshed battery pack and fixture.", + "texts": [ + " First, a modal analysis was carried out to extract all mode shapes and to determine the resonance frequencies between 5 Hz and 2 kHz. Following the modal analysis, a harmonic sine sweep was employed between 5 and 100 Hz, the steps of which are presented in the following section. Lastly, a random vibration profile was applied to the battery block and the response of the structure was obtained. The representative battery pack considered in the study was divided into four main components: the fixture, the bottom plate, the battery cells and the connection elements and the top plate, as shown in Fig. 3. The assembled FEM model of the battery block and the fixture are illustrated in Fig. 4. In the analysis, linear elastic material behaviour was assumed, and the small deflection theory was used. In addition, no nonlinearities were included. The damping coefficients employed in the related simulations were selected based on a previous study [11] by the authors. The FEM model included a total of approximately 15,918 elements and 44,412 nodes. The material properties pertaining to the battery block used in the FEM analysis are given in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001670_icems.2015.7385129-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001670_icems.2015.7385129-Figure2-1.png", + "caption": "Fig. 2. Three-dimensional schematic of a cylindrical rolling-element bearing", + "texts": [], + "surrounding_texts": [ + "Rolling-element bearings are comprised of the following components [10]: \u2022 The outer ring with a slot in its inner surface (inner Local defect or abrasion causes periodic bumps on the motor vibration signal. Amplitude and period of repetition of the bumps depend on the shaft speed, location and severity of defects and characteristic dimensions of the bearing. Depending on the faulty component of the bearing, the bumps frequencies are calculated using (1) to (4) [10]. The vibration frequency caused by local defects on the bearing cage (fc) is calculated as follows: )cos1( 2 \u03b2 c br c D Df f \u2212= (1) Vibration frequency caused by local defect on a ball (fb) is twice its frequency of rotation around itself and is calculated as follows: )cos1( 2 2 2 \u03b2 c b r b c b D D f D D f \u2212= (2) Local defects on the outer raceway and inner raceway cause vibrations whose frequencies (fo and fi) are calculated using (3) and (4) respectively: )cos1( 2 \u03b2 c b r b o D D f N f \u2212= (3) )cos1( 2 \u03b2 c b r b i D D f N f += (4) In the above equations, fr is the mechanical speed of the rotor in r/sec, Nb is the number of balls, Db is the balls diameter, Dc is the average diameter of the outer and inner rings and \u03b2 is the contact angle of the balls (see Fig. 1). Any vibration produced by any bearing defect creates harmonic components in the stator line current whose frequencies are determined as follows: || charactersbf kfff \u00b1= (5) where fs is the stator supply frequency, k=1, 2, 3, ... and fcharacter is the existing vibration frequency. The stator line current harmonics provide non-invasive approach to detect incipient bearing faults [6]. Bearing fault diagnosis studies have generally been carried out using experimental results which offer little flexibility. This paper proposes a method for modeling and simulation of localized bearing faults in squirrel cage induction motors using multiple coupled circuit modeling." + ] + }, + { + "image_filename": "designv11_34_0002056_1350650116631453-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002056_1350650116631453-Figure3-1.png", + "caption": "Figure 3. Forces acting on the roller in the unloaded zone.", + "texts": [], + "surrounding_texts": [ + "According to load balance (Figure 1), force balance equation is written as25,26 W \u00bc Q0 \u00fe 2 X Q cos \u00f01\u00de at UNIV CALIFORNIA SAN DIEGO on February 24, 2016pij.sagepub.comDownloaded from whereW is bearing load, Q0 are the maximum load on roller ( \u00bc 0), is the position angle, and Q is the load distribution at the different angle position. According to conformability of deformation, the deformation can be described as25 \u00bc max cos \u00f02\u00de In accordance with the relationship between the deformation and the contact load, one equation can be attained as25 Q Q0 \u00bc max 1=t \u00f03\u00de where t is the coefficient, t\u00bc 3/2 for ball bearings, t\u00bc 10/9 for roller bearings. Equations (2) and (3) are substituted into equation (1) which can be expressed as follows Q0 \u00bc W ZJ J \u00bc 1\u00fe2 P cost Z ( \u00f04\u00de where 1=J is a constant correlating with the roller numbers Z (see Houpert27 and Wan25). The loading conditions in the loaded zone and unloaded zone are given in Figure 2 and 3, respectively. Qij is the normal force between the roller and inner race, Qoj represents the normal force between the roller and outer race. Taking 1=J \u00bc 4:6, the normal force between the maximum loaded roller and inner race is given by27,25 Qi0 \u00bc 4:6W Z \u00f05\u00de The normal contact force between the maximum loaded roller and outer race is described as Qo0 \u00bc Qi0 \u00fe F! \u00f06\u00de where F! is the centrifugal force of the roller, which is written as F! \u00bc mRm! 2 c \u00f07\u00de where m is the mass of roller. When bearing is operating under lubrication condition, hydrodynamic pressure cannot be ignored. The dimensionless transverse component of hydrodynamic pressure force acting on the roller which is developed by Dowson and Higginson26 using a semiempirical relationship. According to Dowson\u2019s EHL theory, Pij is given by2,25,26 Pij \u00bc 18:4\u00f01 \u00deG 0:3 U0:7 ij \u00f08\u00de where is described as \u00bc Rr=Rm, G is the material parameter and expressed as G \u00bc E0, where represents the viscosity\u2013pressure coefficient of lubricant and E0 represents the equivalent elastic modulus. Moreover, Uij is the dimensionless velocity parameter between the roller and inner race, which is at UNIV CALIFORNIA SAN DIEGO on February 24, 2016pij.sagepub.comDownloaded from expressed as Uij \u00bc 0Uij E0Ri \u00f09\u00de where 0 is the dynamic viscosity of lubricant oil at room pressure, Uij is the entrainment velocity of lubricating oil between the roller and inner race, it is described as Uij \u00bc 1 2 \u00f0Rm Rr\u00de\u00f0! !c\u00de \u00fe Rr!j \u00f010\u00de Poj represents the dimensionless transverse component of hydrodynamic pressure force acting on the roller which is given by2,25,26 Poj \u00bc 18:4\u00f01 \u00deG 0:3 U0:7 oj \u00f011\u00de where Uoj is the dimensionless velocity parameter between the roller and inner race, which is expressed as Uoj \u00bc 0Uoj E0Ri \u00f012\u00de where Uoj is the entrainment velocity of lubricating oil between the roller and outer race, it is described as Uej \u00bc 1 2 \u00f0Rm \u00fe Rb\u00de!c \u00fe 1 2 Rr!j \u00f013\u00de When bearings operate, the roller is driven by friction force between the roller and races. The dimensionless friction force (or traction force) can be described as2,25,26 Fij \u00bc 9:2G 0:3U0:7 ij \u00fe VijIij Q0:13 ij 1:6G0:6U0:7 ij \u00f014\u00de where Vij is the dimensionless velocity parameter between the roller and inner race, which is expressed as Vij \u00bc 0Vij E0Ri \u00f015\u00de where Vij is the relative slip velocity between the roller and inner race, it is described as Vij \u00bc \u00f0Rm Rr\u00de\u00f0! !c\u00de Rr!j \u00f016\u00de where Iij is a dimensionless integral which is a component of sliding friction. The sliding friction force (or traction force) is an integral of viscosity which is a function of contact pressure to describe the tractive effect owing to increased viscosity according to the contact pressure distribution in the contact region. The contact pressure distribution between the roller and the races is well closed to Hertz pressure force distribution. A function of Hertz contact pressure is integrated by contact width, then the traction force due to sliding can be calculated. The dimensionless integral Iij is given as2,25,26 Iij \u00bc 2 Z 4 qij 0 exp G qij\u00bd1 \u00f0 x 4 qij \u00de 2 1=2 d x \u00f017\u00de where qij \u00bc ffiffiffiffi Qij 2 q , and Qij is the dimensionless normal contact force between the roller and inner race which is written as Qij \u00bc Qij lE0Ri \u00f018\u00de Foj is the dimensionless friction force between the roller and outer race, it is given by2,25,26 Foj \u00bc 9:2G 0:3Uoj \u00fe VojIoj Q0:13 oj 1:6G0:6U0:7 oj \u00f019\u00de where Voj is dimensionless velocity parameter between the roller and outer race, which is expressed as Voj \u00bc 0Voj E0Ro \u00f020\u00de where Voj is the relative slip velocity between the roller and outer race, it is described as Voj \u00bc \u00f0Rm \u00fe Rr\u00de!c Rb!j \u00f021\u00de where Ioj is the dimensionless integration, which is given by2,25,26 Ioj \u00bc 2 Z 4 qoj 0 exp G qoj\u00bd1 \u00f0 x 4 qoj \u00de 2 1=2 d x \u00f022\u00de where qoj \u00bc ffiffiffiffiffi Qoj 2 q , and Qij is the dimensionless normal contact force between the roller and the outer race which is given by Qoj \u00bc Qoj lE0Ro \u00f023\u00de It is assumed that the normal roller-cage force Fd of each roller in the loaded zone is equal to each other. Similarly, the normal roller-cage force Fdu in the unloaded zone of each roller is also equal to each other. Fd and Fdu are the dimensionless normal force which are described as Fd \u00bc Fd lE0Ro \u00f024\u00de Fdu \u00bc Fdu lE0Ro \u00f025\u00de at UNIV CALIFORNIA SAN DIEGO on February 24, 2016pij.sagepub.comDownloaded from fdj represents the dimensionless friction force between roller and cage. In the loaded zone, it is expressed as fdj \u00bc f Fd \u00f026\u00de In the unloaded zone, it is expressed as fdu \u00bc f Fdu \u00f027\u00de" + ] + }, + { + "image_filename": "designv11_34_0000329_b978-0-12-397945-2.00009-3-Figure9.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000329_b978-0-12-397945-2.00009-3-Figure9.3-1.png", + "caption": "Figure 9.3 Relationship between the bend angle (y) and the net amount of microtubule sliding (L).", + "texts": [ + " From the recorded images, wave parameters, such as wavelength (l), bend angle (y), and shear angle (s), shown in Fig. 9.2, are obtained. The wavelength is defined as a length consisting of a pair of bends (Fig. 9.2A): in the flagellum of sea urchin sperm, asymmetrical bending waves, consisting of a principal bend (P) and a reverse bend (R), the former larger than the latter, are alternately formed (Gibbons & Gibbons, 1972) (Fig. 9.2B). The bend angle is a parameter proportional to the net amount of microtubule sliding that has occurred at the bending region in question (Fig. 9.3), assuming that there is no twisting of the axoneme or longitudinal compliance of the microtubules (Brokaw, 1991). In the case of asymmetrically beating flagella, an average bend angle of the P and R bends are usually used as a useful parameter. For further analysis of the microtubule sliding in beating flagellum, the shear angle curves showing microtubule sliding as a function of position along the length of a flagellum are useful; they are obtained as a difference in angular orientation between the locus on the flagellum and the basal end of the flagellum (or sperm head), where there is assumed to be no sliding (Gibbons, 1981; Satir, 1968; Warner & Satir, 1974) (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002503_epe.2016.7695684-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002503_epe.2016.7695684-Figure5-1.png", + "caption": "Fig. 5: Scheme of bearing coupling and acting forces", + "texts": [ + " The last vector on the right hand side in (2) is given as a sum nonlinear vectors of bearing forces mentioned above. Nonlinear bearing radial forces determination The approach to bearing modeling used here is based on the methodology presented in [17], but with some simplifications because the main goal is to qualitatively analyze the rotor behavior and to determine dynamic bearing forces. The simplifications consist in the assumption when the rolling elements do not move around the bearing ring circumference and that the radial force transmits one rolling element only (e.g. point in Fig. 5). Further, the housing is supposed to be rigid and it means some modification of the approach used in [17]. The bearing model includes nonlinear contact forces which depends on the shaft center deflection in the bearing cage and on the radial direction which the force acts. Let us express the radial force of the i-th bearing (i=L,R) as follows , (3) where is a unit vector of main radial bearing deformation in the plane perpendicular to rotor axial direction, is nonlinear bearing force which depends on mean bearing stiffness , gives main radial deformation of i-th bearing and is radial bearing clearance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003203_j.est.2020.101417-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003203_j.est.2020.101417-Figure1-1.png", + "caption": "Fig. 1. Schematic of the battery pack.", + "texts": [ + " In the experimental work, it was observed that the battery-pack retained its structural integrity without experiencing any kind of mechanical failures. It was also observed that the outcomes of the finite element simulations are reasonably consistent with the test results. Lithium-ion battery packs have been frequently used in space applications, such as satellites, because of their superior performance and lightweight structures. Due to the challenging flight conditions encountered, especially in the course of orbiting, strict procedures are to be considered in their design and manufacturing phases. Battery packs (Fig. 1) used in satellite missions experience harsh mechanical loads, such as random vibrations and shocks during the launch, climb and stationing on the determined orbit. To meet these demanding conditions, the mechanical structure of the battery packs should be designed and tested according to specific standards, the reliability of which has already been proved. Design, manufacture and test procedures of Li-ion battery packs have already been investigated in recent studies [1\u20134]. While many of these studies [1\u20133] focused on the application of Li-ion battery packs in electric and hybrid electric vehicles, Bolandi [4] and his co-workers approached the design application of Li-ion battery packs in spacecraft vehicles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003713_iecon43393.2020.9254216-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003713_iecon43393.2020.9254216-Figure4-1.png", + "caption": "Fig. 4. Volar plate with the collateral ligaments", + "texts": [ + " The prosthetic hand contains all joins and ligaments as the real human hand. The ligaments and tendons are produced with rubber adapted to each finger join individually. The volar plate and collateral ligaments are developed by rubber with a hardness of 7kg/m and thickness of 0,7mm. Each join is able to realize the flexion/extension with respect to the frontal plane. The thumb abduction/adduction movement is realized in the sagittal plane. For the extensor hood a rubber band of hardness of 2 kg and 0,5 mm thin has been used. Fig.4 shows the assembled structure of a finger with the collateral ligaments, volar plate and extensor hood. The extensor hood is made from rubber with hardness of 5kg/m. This makes it possible to use the servo and step motors to realize the mechanical flexion/extension. The movements of the prosthetic hand are performed by the servo and step motors. Each finger is dotted with one step motor for flexion/extension. The thumb has one servo motor for abduction/adduction. The servo provide an independent motion, which means that each prosthetic hand finger can move independently" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002656_icamechs.2016.7813415-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002656_icamechs.2016.7813415-Figure8-1.png", + "caption": "Fig. 8. Outline of attitude control system", + "texts": [ + " The received signal controls the driving unit and the rotation of each system through a motor driver. Mowing system behave according to the data of the acceleration sensor at a slope. The mowing system can be operated by using only one controller. Therefore, a single operator can operate the system with ease. When making this system moves at a slope, the system is slippery on the slope underside by influence of gravity. First we consider system\u2019s attitude control using an acceleration sensor by this research. Fig. 8 shows outline of attitude control system. It's measured what kind of attitude a system faces to in a slope. And the system corrects an error with the reference attitude by left and right crawler. In particular, the system measures 2 axes of inclination of movement using an acceleration sensor. Then, the system recognizes whether the system is downward-sloping. Finally, velocity differential is given to left and right crawler by attitude correction movement is performed. In addition, slipping at a slope is restrained by attaching the spike wheels in the side of the crawler belt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003357_ilt-04-2020-0152-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003357_ilt-04-2020-0152-Figure1-1.png", + "caption": "Figure 1 Dynamic model of slide-crank mechanism and its impacting force on the crankpin bearing", + "texts": [ + " The dimensions of micro-dimples are then optimized based on the MOGA to ameliorate the ELF. The simulation result of optimized micro-dimples is compared with that of the optimal CB dimensions via the indexes of decreasing the solid contact force, friction force and friction coefficient between the crankpin and bearing surfaces; and increasing the oil film pressure. The CB including a slip crankpin surface moves on a non-slip bearing surface of the engine with the angular velocity v , as seen in Figure 1(a). The impacting force W0 of slider-crank mechanism can influence CB\u2019s lubrication and friction efficiency. Based on the motion equation of slider-crank mechanism and experimental result of the combustion gas pressure at 2,000 rpm (Nguyen et al., 2020; Zhao et al., 2016), the result ofW0 is calculated and plotted in Figure 1(b). Under the impact of W0 changing in both direction and intensity, the load bearing capacity (Wb) of CB generated by the load bearing capacity of the oil film pressure (Wof) and asperity contact (Wac) in the mixed lubrication region must be equal toW0. Thus, the oil film thickness and its pressure (h and p) are changed, which affects the lubrication and friction in CB (Wang et al., 2011). A CB\u2019s hydrodynamic model supported by the W0 and rotated with v inside the bearing was established as in Figure 2(a), and its simple model was also given to its Cartesian counterpart as in Figure 2(b) to optimize the CB\u2019s dimensions (Nguyen et al", + " To calculate the initial values of J1 and J2 in equation (12), an algorithm program developed in the Matlab software is applied to solve equations (2)\u2013(4), (6) and (9)\u2013(11). Also, to simplify the calculation of micro-dimples and lubricant film distribution of CB, the matrix of C is used by the square matrix of x \u00bc y \u00bc 120: The matrix of micro-dimples are also assumed to be uniformly distributed on the bearing surface with n \u00bc m \u00bc 6; and length/width of bearing is Lx \u00bc 2p rb andLy \u00bc B: The data of W0 in Figure 1(b) and original parameters of CB and micro-dimples in Table 1 are used to simulate the results. The calculation process is schematized with the basicmodel in Figure 4. The simulation results of the initial distribution of microdimples and oil film pressures on the bearing surface are obtained and plotted in Figures 5(a) and 7(a), and initial values of J1 and J2 are also obtained. The dimension of micro-dimples is then optimized via the MOGA to ameliorate the ELF. Based on the optimal model in the same Figure 4, the initial population of 240 in 1000 generation with the error 60:01 of fitness values is set for the optimal process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000572_1.4929777-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000572_1.4929777-Figure1-1.png", + "caption": "FIG. 1. (a) and (b) Formation schematics and (c) and (d) SEM images of a free-standing InGaAs/GaAs microtube directly fabricated on Si. The broken white lines in (c) show the borders of U-shaped mesa. The solid red lines in (d) show the ideal border of the arm and central part of as-fabricated microtube.", + "texts": [ + ", reduce the threading dislocation density and surface roughness), 2 lm-thick GaAs buffer was grown on Si (100) substrate by previously reported three-step growth method with thermal cycle annealing (TCA) process.24 Next, 50 nm-thick AlAs sacrificial layer was deposited, followed by a 15 nm-thick In0.23Ga0.77As and a 30 nm-thick GaAs layer as top bilayer. The identical epitaxial structure was also grown on GaAs (100) substrate for comparisons. Illustrated in Figs. 1(a) and 1(b) is the formation mechanism of Si-based free-standing InGaAs/GaAs microtubes. First, U-shaped mesas without any corrugation were defined by traditional photolithography. The size details of the mesa are illustrated in Fig. 1(a). H2SO4:H2O2:H2O (1:8:240) solution was then used to accurately etch into InGaAs layer, which transferred the pattern into bilayers. Thereafter, a starting edge was defined by the secondary photolithography and following etching through AlAs sacrificial layer, making only one side of AlAs layer exposed. Subsequently, highly selective HF:H2O (1:40) solution started to laterally remove AlAs, which rolled up the bilayer from one specific direction (i.e., [010] or [001]) due to strain relaxation. After a certain lateral distance which is determined by a timed etching, only the side pieces of U-shaped mesa continued rolling and formed the rolled-up arms", + " The laser beam spots of approximately 2 lm in diameter were focused on the sample by an optical microscope, and the laser power on the sample was about 2 mW. Scanning electron microscopy (SEM) images of asfabricated Si-based single free-standing InGaAs/GaAs microtube are shown in Figs. 1(c) and 1(d). The outer diameter of the central part of the rolled-up tube (i.e., outer diameter of free-standing tube) is 5.4 lm with 3 rotations and wall thickness of 135 nm, corresponding to the inner diameter of 5.1 lm. The central part of tube is separated from substrate of 270 nm through the supporting from the two arms. As obviously shown in Fig. 1(d), the surface of Sibased free-standing InGaAs/GaAs microtube was extremely smooth, which can be convinced that the structural properties of our Si-based III-V microtubes are as good as their GaAsbased counterparts. Realizing strong room-temperature photoluminescence of III-V semiconductors on Si substrates is invariably a great technological challenge. In general, III-V semiconductors metamorphically grown on Si will exhibit large surface roughness value and high threading dislocation density, which consequently leads to a very serious deterioration in PL properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003295_s00170-020-05693-0-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003295_s00170-020-05693-0-Figure5-1.png", + "caption": "Fig. 5 Illustration of the grinding processes", + "texts": [ + " In this work, to calculate the wheel position and orientation, the conical end-mill was discretized into a finite of segments shown in Fig. 4. Each step of grinding could be viewed as a cylindrical end-mill with a radius rTi. To this end, the wheel path for conical flank grinding could be simplified as solving a series of the flank grinding for cylindrical end-mill. For the cylindrical end-mills, to determine the wheel\u2019s position, a grinding point is introduced. The grinding point is located by the parameter HGPi shown in the broken view of Fig. 5. Based on the grinding point, the wheel\u2019s position [yi zi] can be represented in Eq. (9): yi \u00bc rTi\u2212HGPi sin\u03b2i \u00fe R cos\u03b2i zi \u00bc \u2212HGPi cos\u03b2i \u00fe R sin\u03b2i : \u00f09\u00de Besides, the cylindrical flank grinding can be expressed as Eq. (10) by updating Eq. (7): TW h; \u03b8; t\u00f0 \u00de\u00bc R cos\u03b8 cos \u03c9i t\u00f0 \u00de\u2212R sin\u03b8 cos\u03b2i sin \u03c9i t\u00f0 \u00de \u00fe h sin\u03b2i sin \u03c9i t\u00f0 \u00de\u2212yt sin \u03c9i t\u00f0 \u00de R cos\u03b8 sin \u03c9i t\u00f0 \u00de \u00fe R sin\u03b8 cos\u03b2i cos \u03c9i t\u00f0 \u00de\u2212h sin\u03b2i cos \u03c9i t\u00f0 \u00de \u00fe yt cos \u03c9i t\u00f0 \u00de R sin\u03b8 sin\u03b2i \u00fe h cos\u03b2i \u00fe zi \u00fe v t 2 4 3 5: \u00f010\u00de In grinding processes, the grinding wheel contacts the flank at a curve, which is called the contact curve at this moment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure11-1.png", + "caption": "Fig. 11 Cutting region when peripheral cutting edge does not take part in machining and yB>R1", + "texts": [ + " Integral limits are, a \u00bc arccos yA r b \u00bc arccos yB r 8<: c \u00bc arccos yB r d \u00bc arccos yB r 8<: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f032\u00de (3) For IEi in part III, where ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2\u2212 h\u2212Rc\u00f0 \u00de2 q < yB and R1 <>: c \u00bc \u2212arccos yB r d \u00bc arccos yB r 8<: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f033\u00de (4) For IEi in part IV, where ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2\u2212 h\u2212Rc\u00f0 \u00de2 q > yB, and R1 <>: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f034\u00de (5) For IEi in part V, where yC<>: c 0 \u00bc arccos yC r d 0 \u00bc arccos yB r 8<: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f035\u00de (6) For IEi in part VI, where R2R1, the cutting region also can be divided into six different parts, as shown in Fig. 11. (1) For IEi whose cutting path is involved in part I, where yA<>: \u00f037\u00de (3) For IEi in part III, where yB<>: c \u00bc arccos yB r d \u00bc arccos yC r 8<: c 0 \u00bc arccos yC r d 0 \u00bc arccos yB r 8<: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f038\u00de ( 5 ) F o r I E i i n p a r t V, w h e r e ffiffiffiffiffiffiffiffiffiffiffiffiffiffi h\u2212Rc\u00f0 \u00dep 2 \u00fe yB 2 < r < R2; si \u00bc s2 \u00fe s2 0 \u00fe s1 c \u00bc arcsin h\u2212Rc r d \u00bc arccos yC r 8><>: c 0 \u00bc arccos yC r d 0 \u00bc arccos yB r 8<: a \u00bc arccos yB r b \u00bc arccos yA r 8<: \u00f039\u00de (6) For IEi in part VI, where R2R1, the cutting region can be divided into four different parts, as shown in Fig. 13. In this case, part I, part II, and part III are all the same as they are in Fig. 11. So the corresponding swept area si of the three parts can be calculated by Eqs. (31), (37), and (33). While part IV can be treated the same as part IV in Fig. 10, we can use Eq. (34). Based on the analysis above, swept area si of any IEi on the cutting edge can be calculated according to Eqs. (31) to (39). Although cutter wear cannot be avoided in turn-milling, it can be balanced on whole cutting edge if the cutting parameters are optimized, thereby prolong cutter life and reduce the cost on cutter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002599_s1068366616040024-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002599_s1068366616040024-Figure2-1.png", + "caption": "Fig. 2. Composite inhomogeneous porous bearing.", + "texts": [ + " This paper is devoted to the development of the model of damper gain with the compressible oil layer and porous insert with the variable permittivity under the conditions of the incomplete filling of the clearance with the lubricant material assuming the feed of the lubricant material at radial and circumferential directions. The aim of this work is to generate the specified calculation model of the damper with porous sintered insert and compressible oil layer, which provides the assumption of the simultaneous effect of the set of variable structural and operational factors on the gain. STATEMENT OF PROBLEM The scheme of the damper with a compressible oil film and a porous cage is given in Fig. 1. The equation of motion of rotor in directions r and t (Fig. 2) for nonstationary motion of the axle center can be represented as follows: (1) (2) ( ) ( )[ ] ( ) ( ) 22 2 2 sin cos cos , r Y Y X X dd em e F W K Y dtdt K X u t \u23a1 \u23a4\u03d5\u2212 = \u2212 \u2212 + \u03b4 \u03d5\u23a2 \u23a5 \u23a3 \u23a6 \u2212 + \u03b4 \u03d5 + \u03c9 \u03d5 \u2212 \u03c9 ( ) ( ) ( )[ ] ( ) ( ) 2 2 2 2 cos sin sin . t Y Y X X d ddem e dt dtdt F W K Y K X u t \u23a1 \u23a4\u03d5 \u03d5+\u23a2 \u23a5 \u23a3 \u23a6 = \u2212 \u2212 + \u03b4 \u03d5 + + \u03b4 \u03d5 \u2212 \u03c9 \u03d5 \u2212 \u03c9 396 JOURNAL OF FRICTION AND WEAR Vol. 37 No. 4 2016 AKHVERDIEV et al. Assume the working load W is stationary and oriented according to the scheme in Fig. 2. The load of damper causes initial displacements, which are determined by the relationships and Considering that 0X\u03b4 = .Y YW K\u03b4 = ,X YK K K= = ,e C\u03b5 = ,rT t= \u03c9 and one can represent Eqs. (1) and (2) as follows: (3) (4) The values of forces and are determined by the integration of the pressure in the layer of lubricant material according to parameters r and t. For this purpose, it is necessary to first solve the equation for the pressures in porous ring and the layer of liquid oil, as well as to bring these solutions into accordance by the interface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003108_1350650120925363-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003108_1350650120925363-Figure3-1.png", + "caption": "Figure 3. Coordinate system and solution zone of novel GFCB.", + "texts": [ + " To verify the practical application feasibility of the novel bearing proposed in this article, a pair of GFCBs was used to replace the traditional radial-thrust foil bearing supporting scheme of an air compressor, as shown in Figure 2(b). And two vibration sensors were placed at the position of front and rear conical bearings respectively to monitor the vibration of the air compressor during the actual operation, which could be used to judge the actual supporting effect of GFCB. In the conical coordinate system shown in Figure 3, the Reynolds equation for the novel GFCB is @ @ h3 @p2 @ \u00fe r sin @ @r r sin h3 @p2 @r \u00bc 12u!r2 sin2 @ ph\u00f0 \u00de @ \u00fe 24ur2 sin2 @ ph\u00f0 \u00de @t \u00f01\u00de where and r are circumferential and radial coordinate, respectively, ! is the rotating angular velocity, p is the film pressure, and h is the film thickness, is the half cone angle, u is the gas viscosity, and t is the time. The cone can be expanded into a sector with the following equations R1 \u00bc r1 sin ,R2 \u00bc r2 sin a,L \u00bc \u00f0R2 R1\u00de cos a where R1, R2 are the radius of small end and large end, respectively, r1, r2 are the radius of the expanded sector, and L is the bearing length", + " For the GFCB, the following dimensionless Reynolds equation can be defined @ @ h3 @ p2 @ \u00fe r sin @ @ r r h3 sin @ p2 @ r \u00bc 2 r2 sin2 @ \u00f0 p h\u00de @ \u00fe 2 r2 sin2 @ \u00f0 p h\u00de @ t \u00f03\u00de The gas film thickness is h \u00bc h0 e1 cos cos \u00fe e2 sin \u00fe wd \u00f04\u00de where e1 and e2 are the radial and axial eccentricity, respectively, and wd is the deflection of the support structure. Then the dimensionless film thickness is h \u00bc 1 \"1 cos cos \u00fe \"2 sin \u00fe wd \u00f05\u00de where \"1 and \"2 are the radial and axial eccentricity ratio, respectively, \"1 \u00bc e1=h0, \"2 \u00bc e2=h0, and wd is the dimensionless top foil deflection. To simplify the calculation and improve the accuracy, the threedimensional conical solution domain is transformed to a two-dimensional one through the conformal transformation method, where the coordinate , r\u00f0 \u00de is changed as , \u00f0 \u00de (shown in Figure 3). Then equation (3) can be expressed as @ @ h3 @ p2 @ \u00fe @ @ h3 @ p2 @ \u00bc 2 sin2 @ \u00f0 p h\u00de @ \u00fe 2 2 sin2 @ \u00f0 p h\u00de @ t \u00f06\u00de where \u00bc ln r= sin , d \u00bc d r= r sin \u00f0 \u00de, and \u00bc e2 sin . As a two-dimensional standard rectangle solution domain is applied, the thick plate element model can be used for the top foil deformation solution.20 The bump foil is regarded as a linear spring with the spacing distribution connected to the top foil (Figure 4), and the stiffness matrix k\u00bd e of plate element is of the form k\u00bd e\u00bc kf \u00fe kS \u00f07\u00de For the isotropic case, the stiffness matrices kf and kS due to flexure and shear, respectively are detailed in the reference" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003912_ssci47803.2020.9308340-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003912_ssci47803.2020.9308340-Figure2-1.png", + "caption": "Fig. 2: The modeling of multiple sub-paths; (a) UAV examines a certain range around its position with a predetermined radius, (b) Representation of the straight line perpendicular to the current and next waypoints to determine multiple alternative sub-paths.", + "texts": [ + " The objective function used by DEA* is shown in Equation 2. Towards this aim, the algorithm is widely favorable for reducing computation time especially for realtime path planning in cluttered areas. min Fobj = M\u2211 i=1 fi(v) (2) Once the previous procedure is complete and the global path is obtained, multiple sub-paths are generated around the global path. It is assumed that UAV scans a certain range around its position with a predetermined radius r that the center is the current location of the UAV (See Fig. 2 (a)). It is the same way that the lasers do in a real robot. Then, based on the predefined radius, the obtained global path is divided into equal segments by waypoints. In this way, the distance between two waypoints is under the control of UAV. Now, the main question is how we can determine these alternative subpaths between two pairs of waypoints. To do this, we consider the straight line perpendicular to the next waypoint. We take into account several feasible points upon this line based on the uniform distribution. Then, A* algorithm is applied to find the shortest sub-paths between the current position and each traversable point. Thereafter, all the obtained multiple sub-paths are arranged according to the safest (maximum distance from the moving obstacle) and the shortest length with the global path, and they are stored in a temporary archive (See Algorithm 2). This procedure is repeated for all waypoints. For instance, as can be seen in Fig. 2 (b), from waypoint i to waypoint k, all the possible sub-paths have been determined. Thus, we have a connected graph of all conceivable movements. The local path planning is used after the UAV take-off to follow the determined global path. The idea behind this approach is that it must ensure that the UAV avoids collisions with any moving obstacles during its flight. Hence, considering several alternative sub-paths can be helpful to reach this goal. The moving obstacles can be detected via UAV by measuring the distance from the current position of the UAV and the moving obstacle just in front" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001404_dscc2013-3941-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001404_dscc2013-3941-Figure1-1.png", + "caption": "Figure 1. THE \u201cLANDFISH,\u201d CONSISTING OF TWO LINKS, A NONHOLONOMICALLY CONSTRAINED WHEEL, AND CASTER.", + "texts": [ + " We apply geometric mechanics techniques to establish the equations of motion in terms of the system\u2019s nonholonomic momentum and analyze the system\u2019s equilibrium properties. Finally, we demonstrate its locomotion capabilities under several controllers, including heading and joint velocity control. Nonholonomic vehicles have widely been studied for their ability to locomote solely via interactions with their environment. Many canonical examples are extensions of the Chaplygin sleigh [1], a rigid body supported by a wheel and casters, with the constraint that the wheel cannot slip laterally. This paper introduces the \u201clandfish,\u201d shown in Fig. 1. Like the sleigh, the body has a nonholonomically constrained wheel at the rear, but this system consists of two connected rigid bodies, with the front link being able to rotate freely about the back. The landfish generalizes several Chaplygin sleigh variations in the literature. Osborne and Zenkov [2] studied a sleigh with a moving point mass around the sleigh\u2019s center of mass, but the present system decouples the mass distribution imbalance from the rotational inertia imbalance. Fairchild et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002574_detc2016-60019-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002574_detc2016-60019-Figure15-1.png", + "caption": "Fig. 15 Feature inclusion approach (FIA): including features locally in the CAD environment", + "texts": [ + " The first step in the feature inclusion approach is to identify the contours or layers of the virtual 3D layered prototype for which the feature relation have to be included. Next, from the superimposition of the original model with the layered model, the deviation points between them are calculated, and the points which lie outside the layered model are included within the layered model. The final step is to extrude that contours which correspond to the inclusion of new points. The remaining contours are the previous ones which correspond to the initially sliced model. The schematic of the methodology for a new feature inclusion approach is shown in Fig. 14. Figure 15 shows the results for FIA implementations. The CAD model is locally altered from the sliced contours data and the features loss are prevented through FIA. Therefore, in respect of previous developed global approaches, the proposed technique is more efficient. From the present study it is also observed that there is relation behind the feature loss and facet normal. Moreover, some guideline presented here that can be employed to utilize this knowledge for best-fit pre-processing techniques. If the angle between facet normal and slice direct is 00 or 1800 then there is a possibility to feature loss as flat area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003439_s10854-020-04194-w-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003439_s10854-020-04194-w-Figure7-1.png", + "caption": "Fig. 7 Transformer displacement caused by core magnetostriction", + "texts": [ + "01\u00a0s in a period, it is a special status while the magnetic field and magnetic flow are in symmetrical distribution between left and right limbs. Since the magnetic resistance of core made of silicon steel is far less than magnetic resistance of fixed structure and winding, the magnetic fluxes are concentrated mainly in the core. The measured 1 3 magnetization curves of silicon steel are used and the magnetostriction effect in the transformer numerical model is included by utilizing the magnetostriction curves. Figure\u00a07 shows the transformer deformation distributions of four times (4, 8, 10, 15\u00a0ms) from transient FE simulation when the transformer works with rated frequency 50\u00a0Hz and considering magnetostriction and DC bias. Enlarge all the graphics 1000 times for clear visualization. The deformations of fixed links and transformer shells are obvious and the maximum reached 48.4\u00a0\u03bcm on the fixed rods. It is because the deformation amplified from strain change of core through transformer structures. According to the simulation results, it is obvious that the core corners have the most serious effect with magnetic field circuit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003126_iraset48871.2020.9092240-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003126_iraset48871.2020.9092240-Figure1-1.png", + "caption": "Fig. 1. Mesh curve of uniform stepped rod.", + "texts": [ + " Maximum Torque (M) = 17600 N mm. Modulus of Rigidity (G) = 80 X 103 N/ mm2. Young\u2019s Modulus (E) = 207 X 103 N/mm2. The equation of torsion is [4]: L The equation of torsion strain energy is [4]: IV. ANSYS ANALYSIS Models prepared using CATIA were imported in ANSYS software for analysis. Material properties as mentioned in calculation section were assigned to each model which were meshed (default) later. The mesh generated for the uniform stepped rod, taper rod, and stepped taper rod has been shown in fig. 1, 2 and 3, and their corresponding meshing details have been illustrated in table 2, 3 and 4 respectively. Fixed support was applied at the larger diameter end and torque was applied at smaller diameter end. In all the cases of torsional rod including uniform stepped rod, taper rod and stepped taper rod, the triangular surface meshing was opted to perform finite element analysis. In the case of uniform stepped rod, the total number of elements selected varied between 236 to 412 whereas the total number of nodes used varied between 598 to 904, both of which dependent on the rod geometry (table 2)", + " CONCLUSION The analysis of torsion rod using CATIA and ANSYS was performed to inspect the variation of its strain energy, shear stress induced and total deformation with the change in its shape and cross-sectional area. Three different types of shapes were investigated to obtain an optimized design of torsion rod viz., taper rod, stepped rod (which can be treated as a shaft formed by combining two right cylindrical shafts of different uniform cross-sectional area), and stepped taper rod (which can be treated as a shaft formed by combining two right cylindrical shafts of different tapered crosssectional area). The results obtained by utilizing CATIA and ANSYS are mentioned in fig. 1-9 which are also summarized in table 5 and 6. From the above-mentioned figures and tables, it can be concluded that: 1. For the full length of 110 mm, maximum shear stress was developed in uniform stepped torsional rod with value 2205.3 MPa for L1 = 70 mm, L2 = 40 mm and end diameters D1 = 4.37 mm and D2 = 3.80 mm. For Full length = 130 mm too the maximum shear stress was observed to induce in uniform stepped rod with value 1764.1 MPa and L1 = 75 mm, L2 = 55 mm and end diameters D1 = 4.45 mm and D2 = 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure16-1.png", + "caption": "Figure 16: Unpressurized nonrotating rotor interference CAE analysis with B31 elements for test seals \u2013 VM stress profile at 0.6 mm rotor interference", + "texts": [ + " Rotor rub has been followed by pull back of the rotor to its original position to simulate loading and unloading steps conditions (Figure 14). A displacement contour for an unpressurized-nonrotating rotor interference simulation is detailed in Figure 15 for 0.6 mm radial interference. Due to frictional effects and the cant angle of 45o, the rotor interference deflects the bristles both in the radial and tangential directions. Magnitude of the maximum bristle tip displacement is around 0.825 mm for rotor interference of 0.6 mm. A Von Mises (VM) stress profile at 0.6 mm of rotor interference is given in Figure 16 for the test seals. The maximum VM stress magnitude, which is equal to 86.5 MPa, is observed at the pinch point, as expected. The free-state bristle tip contact force values are simulated for test seals by using rotor-bristle friction coefficients ranging from 0.20 to 0.35. Comparison of simulated BTF values with RTR-unpressurized static BTF measurements are given in Figure 17. As it can be seen from the figure, the analyses results show good agreement with test data for a rotor-bristle friction coefficient range of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000907_cdc.2013.6761101-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000907_cdc.2013.6761101-Figure2-1.png", + "caption": "Fig. 2. Actions of sensor vi Fig. 3. Position of pixel si,l", + "texts": [ + " II. VISUAL SENSOR NETWORKS AND ENVIRONMENT A. Vision Sensors, Environment and Actions In this paper, we consider the situation where n PTZ vision sensors V = {v1, \u00b7 \u00b7 \u00b7 , vn} monitor environment modeled by a collection of m polygons R = {r1, \u00b7 \u00b7 \u00b7 , rm} (Fig. 1). Let the set of position vectors in rj \u2208 R relative to a world frame \u03a3w be denoted by Qj . Suppose that each vi \u2208 V can adjust its horizontal (pan) angle \u03b8i \u2208 \u0398i \u2286 [\u2212\u03c0, \u03c0], vertical (tilt) angle \u03d5i \u2208 \u03a6i \u2286 [0, \u03c0] and focal length \u03bbi \u2208 \u039bi (Fig. 2), where \u0398i, \u03a6i and \u039bi are assumed to be finite sets. In this paper, the notation ai := (\u03b8i, \u03d5i, \u03bbi) \u2208 Ai := \u0398i \u00d7 \u03a6i \u00d7 \u039bi is called an action of sensor vi \u2208 V , and a := (ai)vi\u2208V \u2208 A := A1 \u00d7 \u00b7 \u00b7 \u00b7 \u00d7 An is called a joint action. A collection of actions other than ai is denoted as a\u2212i := (a1, \u00b7 \u00b7 \u00b7 , ai\u22121, ai+1, \u00b7 \u00b7 \u00b7 , an). Once an action ai is fixed, the orientation of sensor vi\u2019s frame \u03a3i relative to \u03a3w and its maximal view angle are uniquely determined [21], which are respectively denoted by Ri(ai) \u2208 SO(3) := {R \u2208 R 3\u00d73| RTR = I3, det(R) = +1} and \u03b2\u0304i(ai)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001027_detc2013-13712-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001027_detc2013-13712-Figure1-1.png", + "caption": "Figure 1. Tilting-pad bearing", + "texts": [ + " However, the nonlinear behavior of rotordynamic system including more complicated bearings and dampers needs to be investigated. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/16/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2013 by ASME Tilting-pad journal bearings have a number of pads, typically four or five. Each pad in the bearing is free to rotate about a pivot and cannot support a moment, which improves the stability margin of rotordynamic systems as compared to fixed pad bearings. A four pad bearing is shown in Fig.1. For a rotor system with tiling-pad bearings, the coupled motion of the shaft and the bearing pads/pivots needs to be taken into consideration [11]. Based on a small perturbation method, often the effect of a tilting-pad bearing is approached by linearized reduced stiffness and damping coefficients (2 by 2) or full dynamic stiffness and damping matrices including all of the pad degrees of freedom [12-13]. In this work, a nonlinear tilting-pad bearing model is presented to describe the behavior of a rotor-bearing system based on a transient analysis to improve accuracy under high vibration conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003571_icra40945.2020.9197009-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003571_icra40945.2020.9197009-Figure2-1.png", + "caption": "Fig. 2. (a) Magnetic system. (b) Dominant coils (labeled in orange color) shift over a control cycle T . The black dashed line separates regions with high and low magnetic field strength. Green circle represents target region. The non-target region consists of four areas (blue color) that are sequentially exposed to low magnetic field. Red arrow indicates fluidic flow direction, and thick black arrow points to an aggregate transiently formed and broken by fluidic flow in the next quarter.", + "texts": [ + " Experimental results confirmed that our proposed swarm control technique is capable of forming aggregates of 1 \u00b5m magnetic particles in a specified region with a mean absolute error of 0.15 mm in positioning the target region and a mean absolute error of 0.30 mm in controlling target region\u2019s radius. Both in vitro (i.e., in a microfluidic phantom) and ex vivo (i.e., in tissue) experiments proved that the formed magnetic aggregates effectively occupied the junctions of blood vessels (or microfluidic channels as phantom) in a physiologically relevant fluidic flow of 80 \u00b5m/s. A system was built to explore 2D aggregation control for embolization. The magnetic system [Fig. 2(a)] consists of four identical magnetic coils. Each coil was inserted with an iron ring core (2 mm thick, 30 mm outer diameter). Electric current supplied to each coil was amplified by a custom-developed amplifier circuit (maximum current: \u00b1 2 A). The magnetic system was integrated with a camera and a translation stage. The sample stage was surrounded by the four magnetic coils and the height of the sample stage is movable along the Z axis for focus adjustment. A syringe pump (Harvard Apparatus Pump 11) was used to control fluidic flow rate", + " To generate aggregates that fully occupy every junction within a specified target region, the magnetic field strength throughout the target region must be sufficiently large to maintain the aggregates\u2019 integrity at the junctions without being broken by fluidic flow. Meanwhile, to prevent aggregates from building up at junctions outside the target region, magnetic field strength in the non-target region must be sufficiently low. Under this low magnetic field strength, the unintended aggregates cannot maintain their integrity under fluidic shear, and they are disassembled into smaller pieces and flushed away. In our control strategy [Fig. 2(b)], in the first quarter of a control cycle (0\u22121/4T ), a pair of adjacent coils function as dominant coils for generating the rotating magnetic field. In 4471 Authorized licensed use limited to: Carleton University. Downloaded on September 19,2020 at 21:50:03 UTC from IEEE Xplore. Restrictions apply. the meanwhile, the other pair of coils act as auxiliary coils to attenuate the magnetic field outside the target region. In the subsequent quarter of the same cycle (1/4T\u22121/2T ), dominant coils and auxiliary coils are shifted clockwise by one coil. In this quarter, the magnetic field distribution changes [see Fig. 2(b)], and the unintended aggregates formed in the previous quarter [e.g., the aggregate labeled by black arrow in Fig. 2(b)], are now located in a region with low magnetic field strength. Dominant coils shift four times to complete a full control cycle T . Throughout a control cycle, the magnetic field in the target region is always sufficiently large, maintaining the integrity of the formed aggregates at junctions of blood vessels. However, outside the target region, aggregates periodically subject to high and low magnetic field strength. They cannot maintain their integrity and are broken into smaller pieces at the junctions. The result is that only the junctions within the specified target region are occupied by aggregates. Within the target region, a number of aggregates are formed; they follow blood flow to the junctions of blood vessels located in the target region and build up into larger aggregates to occupy the junctions [see inset of Fig. 1 and see Fig. 2(b)]. The magnetic field needs to be controlled for generating sufficient magnetic forces among magnetic particles in the target region so that aggregates can maintain their integrity and occupy/stay at the junctions. For an arbitrary position in the workspace, based on the Biot-Savart Law [24], the magnetic field contributed by each coil isBr = \u00b5z 2\u03c0r \u221a (a+r)2+z2 ( a 2+z2+r2 z2+(r\u2212a)2E(k)\u2212K(k))I = RmI Bz = \u00b5 2\u03c0 \u221a (a+r)2+z2 ( a 2\u2212z2\u2212r2 z2+(r\u2212a)2E(k) +K(k))I = ZmI (1) where Rm and Zm are position-related variables determined by the cylindrical coordinate of a target point; m = 1, 2, 3, 4 represents the four coils; a is the radius of the coil; r and z are the cylindrical coordinates [Fig", + "\u2211Ti t=0B(t, p) Ti \u2265 Bcritical, \u2200p \u2208 Q t \u2208 Ti (8) 3) Magnetic field strength for any given position in the non-target region Ui must be lower than Bcritical in every quarter of a control cycle, Ti.\u2211Ti t=0B(t, p) Ti < Bcritical, \u2200p \u2208 Ui t \u2208 Ti (9) where I represents the magnitude of electrical current (in the range of 0 to 2 A in experiments); p represents a point in the workspace; t represents time; and i=1, 2, 3, 4, representing every quarter of a control cycle. The non-target region Ui consists of four regions [Fig. 3(d)] that are sequentially exposed to a magnetic field strength lower than Bcritical in a control cycle. The pair of dominant coils [see Fig. 2(b)] are supplied with current Ii,1(t) and Ii,2(t) and shift by one coil in every quarter of a control cycle. Since the brute-force search gives several sets of current that all fulfill constraints (7)-(9), an optimization function is used to determine the set of current that gives the strongest EXPERIMENTALLY CALIBRATED VALUES Coil 1 Coil 2 Coil 3 Coil 4 a (cm) 2.75 2.70 2.78 2.69 \u00b5 (H/m) 2.96e-3 3.03e-3 2.78e-3 3.43e-3 magnetic field in the target region Q, ensuring that the aggregates in Q maintain their integrity at junctions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001737_j.proeng.2015.12.052-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001737_j.proeng.2015.12.052-Figure2-1.png", + "caption": "Fig. 2. The scheme of the piston-cylinder liner interface.", + "texts": [ + " The function z, is related to the degree of filling of the clearance z, and is characterized by a function that determines the mass content of the liquid phase (oil) in the volume of the clearance between the piston and cylinder using the relationship g11 . For calculating the trajectory of the piston on the lubricating layer in the cylinder, the system of coordinates is fixed to a stationary cylinder. At the start, the origin of this moving coordinate system is at the center of mass (point C) of the moving piston (Fig. 2). According to [5-7], it is assumed that movement of the piston in the cylinder is only in the plane perpendicular to the axis of the piston pin. Given the initial data and the methodology for calculating the trajectory of piston's motion on the lubricating layer, as described in [8-10], we got the dependence of the minimum film thickness minh in function of a crank angle for a diesel engine (Fig. 3). To analyze the process of friction in the contact area we used the discontinuous model where surfaces are represented by asperities of random height (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003565_j.apm.2020.09.018-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003565_j.apm.2020.09.018-Figure2-1.png", + "caption": "Fig. 2. Rotational electromagnetic field and deformations of flexible ring.", + "texts": [ + " When the flexible wheel is deformed, the movable teeth are pushed to move in the radial guide groove of the movable-tooth frame at the same time as meshing with the central wheel. The movable teeth are restrained by the movable tooth frame and the inner tooth profile of the center wheel, and they move along the inner tooth profile of the center wheel, pushing the moving tooth frame to rotate. When the principal axis of the rotating magnetic field is in its initial position, the electric angle is equal to 0 \u00b0 ( Fig. 2 a). When the magnetic field rotates 60 \u00b0 ( Fig. 2 b), the movable teeth move 1/3 of the tooth distance. When the magnetic field rotates 120 \u00b0 ( Fig. 2 c), the movable teeth move 2/3 of the tooth distance. When the magnetic field rotates 180 \u00b0 ( Fig. 2 d), the movable teeth move one tooth distance. Herefore, when the magnetic field rotates 360 \u00b0, the movable teeth move two tooth distances. One rotation of the electromagnetic field corresponds to two tooth distances of the movable- tooth motion; consequently, a reduction ratio occurs, and a large output torque can be obtained. 3. Electromagnetic force and flexible-wheel deformation under eccentricity When an eccentricity occurs in electromagnetic harmonic movable tooth drive system, the initial air gap in the system is unevenly distributed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003711_j.eml.2020.101076-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003711_j.eml.2020.101076-Figure4-1.png", + "caption": "Fig. 4. Completely flexible body model of CBR.", + "texts": [ + " Hence, the initial states of the (n + 1)th single-leg support phase \u0398n+1 can be calculated y a function of the initial states of the nth single-leg support hase \u03d5(\u0398n) as n+1 = \u03d5 (\u0398n) (6) where \u0398n = [\u03b8 (n) 1 \u03b8 (n) 2 \u03b8\u0307 (n) 1 \u03b8\u0307 (n) 2 ] T. Then Eq. (6) is seen as a Poincar\u00e9 map problem. When the CBR is in the stable walking, there exists a fixed point as \u0398n+1 = \u0398n (7) By the use of Eq. (6), the above fixed point (i.e. the initial state) \u0398\u2217 is obtained \u0398\u2217 = \u03d5(\u0398\u2217) (8) The Newton\u2013Raphson method is used to obtain \u0398\u2217 by solving q. (8). .2. Completely flexible body model (CFBM) In order to obtain more accurate walking dynamic responses, e consider the CBR to be a three-dimensional flexible body odel (See Fig. 4). The materials of legs, hip and slope are linear lastic. All system parameters of the CFBM are chosen according o the actual CBR as shown in Fig. 1. The finite element method s applied to discretize the CFBM. To simulate the contact-impact ithin the walking, the penalty function method is adopted. All odes on the undersurface of the footplate are fixed. .2.1. Dynamic equations of CFBM The CFBM investigated in this paper includes contact nonlinarities, but geometric nonlinearities and the material nonlineariies are not considered since the material is pure elastic" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001896_ecce.2014.6953964-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001896_ecce.2014.6953964-Figure18-1.png", + "caption": "Figure 18. Flux density distribution of the prototype IPMSM by FEA", + "texts": [ + " (17) The variation trend of Idh can be measured by locking the rotor, and the estimated d-axis is rotated from \u2013\u03c0/2 to \u03c0/2. The inverter switching frequency is also set to 2 kHz. The measured results of Idh with 100 Hz and 300 Hz injection are illustrated in Fig. 16 and Fig. 17, respectively, which can be analysed as follows. 1) Through changing d-axis and q-axis fundamental current, the machine saliency does not change much, which can be seen from the difference of maximum and minimum value of Idh. This can be explained by the finite element analysis (FEA) simulation as illustrated in Fig. 18. Fig. 18 (a) and (b) show the flux density distribution under no-load and rated-load conditions, respectively. At rated load, the stator is shown to be more saturated whilst the salient rotor saturation remains almost unchanged, i.e. the machine saliency remains when electrical loading increases. This is similar to inset IPM machine [23]. 2) The estimated rotor position corresponding to the maximum Idh shifts when fundamental q-axis current increases, which is due to cross-coupling magnetic saturation effects (the increase of angle \u03b8m in (17))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001408_jae-121631-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001408_jae-121631-Figure2-1.png", + "caption": "Fig. 2. Experimental prototype. Fig. 3. Suspension principle \u2013 (a) rotational angle of magnet is 0 degree; (b) rotational angle of magnet is 20 degree.", + "texts": [ + " Thus the authors also have proposed an improvement method to decrease the zero suspension force of the original magnetic suspension system [10]. However, in this paper, the performance of the improvement method with a special magnet is studied deeply. Some IEM analysis results and measurement experimental results are shown and compared with those of the original magnetic suspension device. Finally, the simulation and experiment results of suspension using the improvement device are shown, and the results indicate that the suspended object can be levitated robustly. The proposed experimental prototype is shown as Fig. 2. This prototype consists mainly of a disk-type permanent magnet, a rotary actuator containing a gear reducer and an encoder, a pair of opposite \u201cF\u201d shape permalloy cores, a suspended object and an eddy current sensor. The magnet that is located in the opposite \u201cF\u201d shape cores is a neodymium magnet and magnetized in radial direction. The diameter of the magnet is 30 mm and the thickness is 10 mm. A rotary actuator behind of the magnet drives the magnet rotate. The actuator that has an encoder measuring the angle of the magnet cannot be seen in Fig. 2. The thickness of the two cores is same as the magnet. The suspended object is installing on a linear rail, and can move in the vertical direction only. The suspension principle can be understood from Fig. 3. This figure shows a schematic diagram of a disk permanent magnet, two opposite F-type iron cores and a suspended object. In order to understand easily, we assume that there is no flux leakage to the air in this magnetic suspension system. Figure 3(a) shows that the magnetic poles of the magnet are aligned in the vertical direction, and the N pole is at the upper side and the S pole is in the lower side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001925_ceidp.2014.6995760-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001925_ceidp.2014.6995760-Figure6-1.png", + "caption": "Fig. 6. Electric field distribution of croSS-aim in phase A", + "texts": [ + "00 @20141EEE 284 increase about 12.9 % to 26.7 % of total voltage, while the voltage undertaken by tension insulators decrease about 0 % to 8.7 % of total voltage. The potential distribution curves taken thorough the center of post insulator and tension insulator in phase A are shown in fig. 5. The potential distribution inside post and tension insulators is basically the same. The curve near flange drops significantly when compared to the middle part, it indicates the high voltage drop near flange. Fig. 6 shows the electric field distribution on the surface of composite cross-arm in phase A. Similar results can be derived from the other two phases. It can be seen that the place near the end of the silicone rubber sheath suffers much higher electric field strength than the middle. Table 2 gives the maximum electric field strength on the surface of sheath and fittings in three phases. It can be seen that the maximum electric field strength on the surface of end fittings and yoke plates is below 2500 Vlmm, it satisfies the critical value" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure17-1.png", + "caption": "FIGURE 17. TEST RIG SET-UP FOR BRUSH SEAL WEAR MEASUREMENT", + "texts": [ + ") Figure 16 shows the impact of a 45% larger fence height to the leakage rate. It was found that the leakage of VBD brush seal is not sensitive to the backplate fence height within steam turbine application range. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/76943/ on 01/29/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 8 Copyright \u00a9 2013 by ASME In order to characterize the endurance of the VBD brush seal to wear, tests were performed on a rotational test rig (Figure 17.) The high-speed rotational test rig, known as the C5R, is capable of spinning up to 2095 rad/s (20K rpm) and can accommodate rotors up to 152.4 mm (6\u201d) in diameter. The rig can be pressurized with either nitrogen (up to 0.35 MPa (50 psi)) or air (up to 3.45 MPa (500 psi)). Details of this test setup are described in [8] and [9]. For this test, sub-scale test seals were fabricated with a 127 mm (5\u201d) bristle pack diameter to fit the rotor size of the C5R. Cost-E and CrMo-V test rotors were machined and T-F-T straight cut, VBD brush seals were installed with a 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002305_b978-0-12-405935-1.00010-1-Figure10.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002305_b978-0-12-405935-1.00010-1-Figure10.1-1.png", + "caption": "FIGURE 10.1 A boundary layer forms when a viscous fluid moves over a solid surface. Only the boundary layer on the top surface of the foil is depicted in the figure and its thickness, d, is greatly exaggerated. Here, UN is the oncoming free-stream velocity and Ue is the velocity at the edge of the boundary layer. The usual boundary layer coordinate system allows the x-axis to coincide with a mildly curved surface so that the y-axis lies in the surface-normal direction.", + "texts": [ + " In addition, boundary-layer phenomena provide explanations for the lift and drag characteristics of bodies of various shapes in high Reynolds number flows, including turbulent flows. In particular, the fluid mechanics of curved sports-ball trajectories is described here. The fundamental assumption of boundary-layer theory is that the layer is thin compared to other length scales such as the length or radius of curvature of the surface on which the boundary layer develops. Across this thin layer, which can exist only in high Reynolds number flows, the velocity varies rapidly enough for viscous effects to be important. This is depicted in Figure 10.1, where the boundary-layer thickness is greatly exaggerated. (On a typical airplane wing, which may have a chord of several meters, the boundary-layer thickness is of order one centimeter.) However, thin viscous layers exist not only next to solid walls but also in the form of jets, wakes, and shear layers if the Reynolds number is sufficiently high. So, to be specific, we shall first consider the boundary layer contiguous to a solid surface, adopting a curving coordinate system that conforms to the surface where x increases along the surface and y increases normal to it" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000872_iweca.2014.6845682-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000872_iweca.2014.6845682-Figure3-1.png", + "caption": "Figure 3. First mode shape of the workpiece.", + "texts": [ + " Furthermore, experimental modal analysis ( EMA) and strain gauges experiment are performed using DASP system and relevant equipment from China Orient Institute of Noise &Vibration. Finite element modal analysis and EMA for single workpiece, VSR equipment and assembly of them are conducted. The natural frequency comparison of order one is listed in table I. The FEA result of the single workpiece is extremely faithful to the EMA. It can be concluded that FEA is reasonable and precise. The natural frequency of the workpiece is found to be quite high. As can be observed in Fig. 3, notable mode shape is revealed in the part of cross thin wall. Therefore, it is the VSR equipment through the vibration platform that is capable of treating the workpiece. The mode shape of the assembly in the FEA is shown in Fig. 4. It should be pointed out that the VSR experimental application is based on the FEA results. The workpiece is placed on the strongly vibratory and easily clamped position, and three rubber supporters are placed on the nodal lines regions. In the FEA results, the natural frequency of the VSR equipment is 65" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003051_s10409-020-00937-4-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003051_s10409-020-00937-4-Figure2-1.png", + "caption": "Fig. 2 a Geometry of the pinned-clamped elastica and b the shape of the pinned-clamped elastica is described by part of a second-ordermode", + "texts": [ + " To obtain the stiffness\u2212curvature curves efficiently, we firstly record typical pinned-clamped elasticas based on the second-order mode and then find pinned-clamped elasticas that correspond to different points on stiffness\u2212curvature curves quickly. The elastica is initially straight and buckles to the first-order mode under external forces at two ends, which induces the end-shortening. After that, two ends are pinned at A1 and A2 as shown in Fig. 1a. The material point C is quasi-statically controlled downward under force Q. Here, we regard the whole elastica as two components A1C and A2C as shown in Fig. 1b. Elasticas A1C and A2C can be regarded as two pinned-clamped elasticas and a general pinned-clamped elastica is shown in Fig. 2a. Then we study the shapes of the whole elastica by the stiffness\u2212curvature curves of two pinned-clamped elasticas, which describe the curvature at the clamped end as slope \u03b2R changes. To obtain the stiffness\u2212curvature curve of a pinned-clamped elastica, shapes of the elastica with different \u03b2R at the clamped end should be determined. Here elastica is inextensible and its weight is neglected. Horizonal location and deflection of the controlled material point are described by x0 and h. \u03b1 is the angle between A1C and A2C ", + " From one second-order mode, we obtain pinned-clamped elasticas with different arc-lengths. This can be realized numerically which is introduced in Sect. 2.3. Shape of one pinned-clamped elastica here can be described by one part of a second-order mode, where the location of the clamped end is on the second-order mode. The rescaling can greatly improve the efficiency of locating stiffness\u2212curvature curves. The stiffness\u2212curvature curve depends on the geometrical features of the pinned-clamped elastica such as L/L0 and L0 in Fig. 2a. Given L/L0 and L0, the stiffness\u2212curvature curve is obtained by varying \u03b2R. So, pinned-clamped elasticas with different L/L0, L0 and \u03b2R are needed for the stiffness\u2212curvature curve. But pinnedclamped elasticas with the same L/L0 but with different L0 have similar stiffness\u2212curvature curves, and thus only one stiffness\u2212curvature curve is needed, while others are obtained by rescaling. An example of rescaling is given as follows. We firstly obtain one pinned-clamped elastica AC in Fig. 2b and then take rescaling to make L0 to be L0 in Fig. 2a. Size of curve AC is L0/L0 times as much as that of AC . At the clamped end, from Eqs. (7) and (8), internal forces Ft and Fn of AC is (L0/L0) 2 times as much as that of AC . From Eq. (6), k of AC is L0/L0 times larger than that of AC . Based on these regulations, we can easily obtain the pinnedclamped elasticas with the same shape with AC . Because one pinned-clamped elastica corresponds to one point on the stiffness\u2212curvature curve, given one stiffness\u2212curvature curve, stiffness\u2212curvature curves of pinned-clamped elasticas with the same L/L0 but with different L0 can be obtained quickly", + " Geometrically, similar transformation of the shapes of individual components can be used to assemble the two components together to form the whole elastica. Similar transformation of a shape is divided into four processes, such as translating, rotating, flipping and rescaling. Translating changes the location and rotating changes the direction, but translating and rotating have no influence on the force and curvature. The magnitudes of parameters after similar transformation are only related to rescaling. Besides rescaling, flipping is also important which inverts component. For example, if strip AC in Fig. 2a, whose main part is convex, is inverted relative to horizonal direction, the main part of the transformed shape becomes concave. Four kinds of shapes are distinguished by whether a component is inverted or not. That either component is inverted can induce two kinds. That both components are inverted or not inverted can induce the remaining two kinds. If elastica A1C or A2C is inverted, \u03b2R and kR change their signs. Shape of each pinned-clamped elastica can be determined by kR and \u03b2R, which corresponds to one point on its stiffness\u2212curvature curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002690_fit.2016.066-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002690_fit.2016.066-Figure4-1.png", + "caption": "Fig. 4. CAD Model from DATCOM", + "texts": [], + "surrounding_texts": [ + "PEM is an established algorithm used to predict and update the estimates of the model parameters. If u(t) is the input and y(t) is output then the past inputs and outputs data can be collected in a vector ZN as given in (10). ( ) ( ) ( ) ( ) ( ) ( ){ } 1 , 1 , 2 , 2 . , (10)NZ u y u y u N y N= \u2026\u2026 On the basis of this past data, the next model output is predicted using (11). ( ) ( )1| \u02c6 (1 )1 1t my t t f Z \u2212\u2212 = Then parameterization of predictor is done in (12), in terms of the finite dimensional parameter vector \u2018\u03b8 \u2019. ( ) ( )1\u02c6 | (1, 2) t m t fy Z\u03b8 \u03b8\u2212= Error can be calculated by comparing the predicted and actual output as given in (13). \u02c6 \u02c6e( ) = y - (t | ) (13)t my\u03b8 \u03b8 Finally parameter is estimated from the model parameterization and the observed data set ZN in a form that the error is minimized in a suitable norm." + ] + }, + { + "image_filename": "designv11_34_0000584_j.jsc.2014.09.031-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000584_j.jsc.2014.09.031-Figure4-1.png", + "caption": "Fig. 4. Tape after two folds.", + "texts": [ + " We could argue that this fold is simply the fold (O1) along line EF. For the sake of the symmetry to be observed in Lemma 2, we use the fold (O5) here. With the tapes shown in Figs. 3(a) and 3(b), the proof is straightforward by elementary geometrical reasoning, and hence is omitted. After the fold by which we constructed the isosceles triangle GFE, we perform another fold to construct another isosceles triangle FHG in such a way that both isosceles triangles overlap, sharing the segment FG as shown in Fig. 4. Lemma 2 expresses the construction involved and the resulting geometrical property. Lemma 2 (Isosceles trapezoid). For all origami ABCD, point E \u2208 AB, point F \u2208 CD and point G \u2208 CD such that O5(G, EA, E, EF), for all point H \u2208 EB and point I such that O5(F, EB, G, GH) \u2227 I \u2208 DEFF \u2227 I \u2208 CGHG, we have |EF| = |GH|, and furthermore |IG| = |IF|. Note that the expression DEFF is parsed as (DEF)F, which denotes the line passing through point F and the reflection of point D across line EF, as explained in Section 2", + " After the successful construction, we try to prove a proposition of the above structure. The clause (ii) is true for some objects including o\u2032 1, . . . , o \u2032 k . Therefore, the proposition is not vacuous. Now let us return to Lemma 2. An intuitive proof of the property that the polygon FEHG is an isosceles trapezoid would be by two applications of Lemma 1. However, this is not completely satisfactory. If point I were not introduced, Lemma 2 would cover one more configuration besides the one shown in Fig. 4. Figs. 5(a) and 5(b) show the other configuration. We would need to show the goal |EF| = |GH| holds for this case, too. We see that this is impossible by observing Fig. 5(a). To eliminate this undesirable configuration, we introduce point I, which is not constructible for the undesirable case. In this case, lines DEFF and CGHG are parallel. Point I will be needed later for Theorem 3 as well. We will postpone the description of our proof until Section 5 since the proof scheme is the same in other constructions. We move on to the construction of a pentagonal knot. On the tape shown in Fig. 4, we perform a fold (O5) along a fold line that passes through F, to superpose H and CG. We have two fold lines, say m1 and m2 to make this possible. The fold along m1 creates the shape as shown in Fig. 6. Point K is created at the intersection of m1 and line CG (in Fig. 4). The other case of the fold (along m2) constructs the shape shown in Fig. 7(b), where point K is not constructible. We now see how to fold the desired knot. The following shows the Eos program to construct the pentagon EHGKF. Program P1 (Construction of a pentagonal knot). 1. BeginOrigami(\u201cPentagonal knot\u201d, {150, 10}); 2. NewPoint({E \u2192 {45, 0}, F \u2192 {43, 10}}); 3. HO(\u2203m,m:Line\u2203n,n:Line\u2203g,g:Point\u2203h,h:Point\u2203i,i:Point(O5(g, EA, E, m) \u2227 m = EF \u2227 O5(F, EB, g, n) \u2227 n = gh \u2227 g \u2208 CD \u2227 h \u2208 AB \u2227 i \u2208 DmF \u2227 i \u2208 Cn g), MarkPointAt \u2192 {G, H, I}, MarkPointOn \u2192 {{CG, K}}, Handle \u2192 {A, B}); 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000914_20130825-4-us-2038.00108-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000914_20130825-4-us-2038.00108-Figure10-1.png", + "caption": "Figure 10. The geometry and notations for the bucket during bucket filling.", + "texts": [], + "surrounding_texts": [ + "While the forces from the weight and inertial components of the machine can be accurately calculated with the presented model from catalog data, the digging force on the bucket is dependent on many factors. Hemami et. al (1994) present a model for this force based on a published formula. However, the force in their formula is a simple, linear function of the cutting edge angle, and the other constants need calibration. A new approach was followed, using force and mass balance analysis in the bucket and in the digging face. Figures 10 and 11 show the geometries and notations for the bucket and the face, respectively, during bucket filling. From fiction in the bucket and lifting of center of mass, the horizontal force is: (21) where (22) Substituting (22) to (21) gives the maximum value for force F: (23) From fiction in the muck pile along friction planes according to , the force balance equation is : (24) In equation (24), V is the volume of the material being sheared and pushed up: (25) The additional force in Eq. (25) is a horizontal force component to lift the amount of mass, expressed from energy balance: ( ) (26) where (27) (28) Substituting (25)-(28) into (24) gives: (29) The force F(x) in equation (23) and (29) are equal under slow, continuous motion, neglecting initial force, and under the condition of a positive F for the bucket. If this condition does not hold, the result will assume tensile force in the muckpile, a physical impossibility. This scenario must be checked before the balance equation is used for solution. After rearrangement, a quadratic equation is obtained for H(x): ( ) , where: (30) ( ) (31) (32) [ ] (33) The vertical distance of the center of mass in the bucket ( is measured from the bucket digging edge. The value of depends on the angle of the digging direction , the bucket angle , and the shape and the instantaneous fill factor, Cm, of the bucket. An approximate formula can be derived for simplified bucket geometry as follows: (34) In Eq. (34), is instantaneous mass of loaded material, M is the mass capacity of the bucket, K is the distance of the line of the center of gravity along which the material moves within the bucket during loading, and L is the opening height of the bucket. The ratio in Eq. (34) m/M is the instantaneous fill factor, Cm: (35) The solution for Eq. (30) provides the height of the material, H, right in front of the bucket loading edge: \u221a (36) Height, H, is a function of several parameters in Eqs. (31)- (34). Two groups of parameters can be defined: Group P1, involving only constant material properties, and group P2, involving digging control and face geometry parameters which are all variable: (37) ( ) (38) The operator of the hybrid machine in \u201cbucket steering\u201d kinematics directly controls and , while is determined by the shape of the fallen face, all variable with the instantaneous positions of the bucket edge. The question is how the bucket filling process plays out from the spotting point at Cm=0, to its completion to Cm=1. In other words, how will Cm(x) depend on (x), (x), and (x)? It must be pointed out that the ultimate hybrid control would be the one which directly controls Cm from 0 to 1 with ease and minimum energy, time and machine wear. In order to connect H to the movement of the bucket along x, another equation is needed. It must be realized that the material slides into the bucket and it does not directly follow its displacement. It is possible to define two elastic parameters, one for the material in the bucket, kB, and one for the pile, kF, as the derivatives of the horizontal forces with respect to x. From Eqs. (23), (29) and (33) we obtain: ( ) (39) (40) The bucket is moved by dx which causes material movements, dxB into the bucket and dxF into the face: (41) With the kB and kF elastic constants, dxB can be expressed: (42) Substitution of Eq. (42) into Eq. (41) and simplifications give: (43) Since , dm can be expressed with the use of Eq. (43): (44) Substituting ycp from Eq. (34) yields, after rearrangement, a differential equation for m as a function of x: (45) Note that m/M=Cm in Eq. (45). The solution for the digging force by moving the bucket is now complete. Starting with a small m and H, first, the dm mass has to be determined for H from the solution of Eq. (45). Next, the new H value can be evaluated from Eq. (36), and the process can continue along the digging trajectory. The digging force F(x), is given by Eq. (29), assuming a positive value for the bucket. If this condition does not hold, the result predicts tensile force in the muckpile, a physical impossibility. This scenario must be checked and corrected by setting the digging depth, H, as a free parameter under direct motion control before Eq. (29) is used for the solution. The procedure was coded into the Hitachi 3500 excavator analysis model in Matlab. The analysis results are shown in Figs. 12 trough 14. The model simulated the motion of the machine along a given trajectory of the bucket edge and variable bucket angle, shown in Fig. 12. Figure 13 shows the joint torques from the lowest to the highest point along the digging part of the trajectory. Figure 14 depicts the cumulative work generated by the joint cylinders assuming 100% energy efficiency and no recuperative energy recovery during lowering of any joints during bucket loading. In the sections of clockwise (lowering) motion, the energy was equated with zero." + ] + }, + { + "image_filename": "designv11_34_0003808_1.g005098-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003808_1.g005098-Figure2-1.png", + "caption": "Fig. 2 Kinematic properties of the twoUAVs,U1 andU2, attached using a rigid link, and the point on the link, U.", + "texts": [ + " The guidance laws for the point-to-platform transfer extension and simulation results are presented in Sec. VI. Finally, some concluding remarks are given in Sec. VII. In this section, first, the kinematics of the 2-UAV system are described, followed by the kinematics-based nonlinear engagement equations between a point on the rigid link and a moving platform. Finally, the control objectives, specific to the cases of stationary and moving platforms, are stated. The kinematic equations of the 2-UAV system are derived, following the schematic shown in Fig. 2, based on the following: 1) The 2UAVs,U1 andU2, are transporting a rigid link of lengthL. 2) The point U on the link is at a distance of \u03c1L from UAV U1, where 0 \u2264 \u03c1 \u2264 1. 3) The reference frames, in the 2-DCartesian space and denoted by X\u2013Y, for each of theUAVs and the pointU on the link are nonrotating and nonaccelerating frames. 4) The 2UAVs and the platform are constrained tomove only in the X\u2013Y plane. The frame defined in this plane, which is used to measure the line-of-sight (LOS) angle and the separation distance, is nonrotating and stationary", + "G 00 50 98 \" aUXeq aUYeq # \" cos\u03b8 \u2212 sin\u03b8 sin\u03b8 cos\u03b8 #\" aUreq aU\u03b8eq # \" cos\u03b8 \u2212 sin\u03b8 sin\u03b8 cos\u03b8 #\" aPr aP\u03b8 # \" cos\u03b8 \u2212 sin\u03b8 sin\u03b8 cos\u03b8 #\" cos\u03b8 sin\u03b8 \u2212 sin\u03b8 cos\u03b8 #\" aPX aPY # \" aPX aPY # (29) Thus, the acceleration components of UAVUequal those of platform P. As a result, it is able to track the platform. It should be highlighted that for a platform moving with constant velocity, that is, where aPX aPY 0, the result shown in Eq. (29) indicates that aUXeq aUYeq 0; that is, once the UAV has aligned itself with the platform and is also moving at the speed of the platform, it does not accelerate in any direction. From Fig. 2, the angle \u03b2L, measured positive counterclockwise about the X axis, denotes the orientation of the rigid link connecting the two UAVs. The rotation of the link is modeled as _\u03b2L \u03c9L; _\u03c9L aL (30) where \u03c9L and aL are the angular velocity and the acceleration of the link, respectively. These states influence the acceleration inputs of the UAVs according to Eq. (2) in Sec. II.A. Consequently, the angular acceleration aL is designed to achieve a desired orientation of the link, \u03b2Ldes, while following the platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003528_j.procir.2020.09.129-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003528_j.procir.2020.09.129-Figure4-1.png", + "caption": "Fig. 4. 3D transport route of Ni element", + "texts": [ + " At the top region, the liquid metal flows backward along the weld pool surface and then forward along the longitudinal section. At the bulging region, the liquid metal flows backward along the boundary of the weld and its flowing direction changes later to become forward. These two dominant circulations are almost separated by the narrowed region. It implies that the formation of the bulging may show significant influence on the basic flow pattern of the weld pool. The transport of the filler material is directly determined by the thermo-fluid flow. Fig. 4 shows the Ni transport in the weld pool. The liquid filler material with higher Ni content first impacts on the keyhole front wall. Subsequently, the filler material flows backward along the lateral side of the weld pool. As the flowing direction of the top circulation changes from backward to forward, the added Ni element is brought to the forepart of the weld pool. This vigorous top circulation produces a homogeneous Ni distribution at the top region. Since the weld pool is significantly narrowed between the top region and the bulging region, the downward channel is blocked" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001011_0954406213484223-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001011_0954406213484223-Figure2-1.png", + "caption": "Figure 2. Dual-shaft SFD test rig: schematic (top), photograph (bottom). SFD: squeeze-film damper.", + "texts": [ + " The unbalance excitation forces in the x, y directions from each rotor can be expressed as Fx1 \u00bc U1 2 1 sin 1t, Fy1 \u00bc U1 2 1 cos 1t \u00f01a; b\u00de Fx2 \u00bc U2 2 2 sin 2t\u00fe \u20192\u00f0 \u00de, Fy2 \u00bc U2 2 2 cos 2t\u00fe \u20192\u00f0 \u00de \u00f02a; b\u00de The initial phase angle \u20192 between the unbalances is arbitrary as they are on separate rotors. It is revealed in this article that the choice of \u20192 is highly influential on the predicted response for certain speed ratios but markedly less so for others. With a few exceptions2,4,8,9 the vast majority of research on nonlinear rotordynamics does not relate to the engine types depicted in Figure 1. Holmes and Dede8 used the same twin-shaft SFD rig of this study (Figure 2), in which the two independently driven unbalanced rotors are nonlinearly coupled through the flexibly mounted housing of an SFD bearing without a retainer spring. This reproduces the effect of the nonlinear dynamic coupling between the rotors in a real aircraft engine, where the casing, of which the housings are part, is flexible (see schematics of Figure 1). Holmes and Dede8 explained the occurrence of combination frequencies using elementary analysis. This rudimentary work was performed at a time when nonlinear analysis techniques of even simpler single-shaft systems were still in their infancy", + " The efficacy of these methods relative to conventional methods was demonstrated in a series of theoretical studies on representative two/three-spool engine models provided by a leading manufacturer1\u20134 (Figure 1). In view of these recent advances in computational techniques, it was considered opportune to recommission and upgrade the dual-shaft SFD test rig originally used by Holmes and Dede,8 for the more comprehensive and detailed experimental and theoretical study that is reported in this article. The test rig is illustrated in Figure 2. An oil annulus (SFD) surrounds one of the rolling element bearings of rotor 1 (the Low Pressure (LP) rotor). J refers to the centre of the SFD journal (which is the outer race of the rolling element bearing and is prevented from spinning). B is the centre of the flexibly mounted SFD housing, which is rigidly attached to the housing of at PRINCETON UNIV LIBRARY on January 13, 2014pic.sagepub.comDownloaded from the adjacent rolling element bearing of rotor 2 (the High Pressure (HP) rotor) through a bell-shaped connector", + " Denoting the HP rotor speed as 2, it is noted that, for a given recorded LP rotor speed 1 and given prescribed nominal speed ratio S, the limitations of the speed controllers and slight fluctuations of the motors meant that the actual recorded value of 2 was not exactly equal to S 1. Hence, in this research, distinction is made between the nominal HP speed S 1 and the actual recorded (\u2018experimental\u2019) HP speed. This distinction is important as far as the theoretical predictions are concerned. The IRM and RHBM are extensively documented by Pham and Bonello.1,2 These computation techniques require the modal parameters of the linear part at zero rotational speed. The linear part is defined as the system resulting when the SFD is replaced by a gap. Hence, in the present case (Figure 2) the linear part comprises two uncoupled subsystems: (a) the pinnedfree rotor 1 and (b) the sprung-pinned rotor 2. The modal parameters of subsystem (a) were calculated by finite element (FE) modelling. The modal parameters of subsystem (b) were obtained by combining the FE model of rotor 2 and bell-shaped connector with the modal parameters of the flexibly-mounted SFD housing (determined experimentally in isolation). The total number of modes R considered in the nonlinear rotordynamic analysis was 10", + " These have been expressed using the same notation and formulation of Pham and Bonello,1,2 enabling a researcher to input these into the formulae of their work1,2 (which cannot be reproduced here for reasons of space). It is noted that the previous applications of the IRM and RHBM involved the simulation of the vibration response of whole-engine models1\u20134 and the modal parameters used could not be published due to confidentiality concerns of the engine manufacturers. Hence, the presentation of the modal input in Appendix 2 in conjunction with the algorithms in Pham and Bonello1,2 serves to elucidate the application of these computational methods. With reference to the system in Figure 2, the IRM and RHBM relate the response at the SFD (J relative to B) to the excitations acting on the linear part as:1,2 the unbalance forces, SFD forces, weight of rotor 1 and the gyroscopic moments on the two rotors. The SFD forces are nonlinear functions of the displacements and velocities of J relative to B in the x, y directions. The x, y relative displacements and/or velocities at the SFD constitute the unknowns of the IRM or RHBM process, and, once determined, the vibration response at all points on the rig is fully defined", + "comDownloaded from S nominal value of rotor speed ratio 2= 1 U1,U2 unbalances on rotors 1 and 2, respectively (kg m) 1, 2 speeds of rotation of rotors 1 (\u2018LP\u2019) and 2 (\u2018HP\u2019), respectively (rad/s) 2f gexp, 2f gnom Experimental and nominal values, respectively, of the speed of rotor 2 (rad/s) \u20192 arbitrary initial phase angle between the rotors 1 and 2, equation (2a, b) (rad) $ fundamental frequency of RHBM solution (\u00bc 1=q) Appendix 2 Computational input details of test rig Table 1 gives the modal parameters of the linear part of Figure 2 at zero rotational speed. !r is the natural frequency of mode r and t r\u00f0 \u00de J , t r\u00f0 \u00de B , etc. are the corresponding mass-normalised eigenvectors evaluated at the relevant degrees of freedom. These degrees of freedom correspond to the directions and locations of the excitations acting on the nonlinear rotating system.1,2 Hence, t r\u00f0 \u00de J defines the modal displacements at J (where the SFD forces act) in mode r and its entries correspond to the x and y directions, respectively (likewise for t r\u00f0 \u00de B ); t r\u00f0 \u00de u 1\u00f0 \u00de defines the displacements in mode r in the x, y directions at the point of application of the unbalance mass on rotor 1 (likewise for t r\u00f0 \u00de u 2\u00f0 \u00de )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001562_sii.2013.6776656-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001562_sii.2013.6776656-Figure1-1.png", + "caption": "Fig. 1. The intelligent walking-aid robot.", + "texts": [ + " In the algorithm both human and robot control the walking aid robot by using reinforcement learning to dynamically adjust the human\u2019s control weight. The intelligent walking aid robot is introduced in section II. The reinforcement learning based walking-aid robot shared control algorithm is given in section III, and experiment results are presented in section IV. II. INTELLIGENT WALKING-AID ROBOT The intelligent walking-aid robot consists of a support frame, a mobile base, an array of sensors, a signal processing system, a rechargeable battery, and a controller. The composition of the mobile base is shown in Fig. 1, a detailed description is shown in our preliminary study [9]. The robot uses a special arrangement of one-dimensional push-pull force sensors instead of the expensive six-axis force/torque sensor most commonly found in such devices to realize the measurement of interactive force and torque between the human and the robot. The force sensors are mounted at each end of the U-shape armrest, measuring the push/pull force of each hand. The armrest is vertically fixed with the support frame by four linear bearings placed horizontally" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000382_icra.2015.7139501-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000382_icra.2015.7139501-Figure1-1.png", + "caption": "Fig. 1. Typical brachytherapy setup including ultrasound probe and perineal template", + "texts": [ + " INTRODUCTION Prostate brachytherapy is a procedure whereby long flexible needles containing radioactive seeds are inserted under ultrasound guidance into the prostate to treat cancerous tissue. The success of this surgery depends on the accuracy with which the seeds are placed within the prostate with respect to pre-planned target locations [1]. During insertion, however, each needle will bend away from its target location, requiring continuous observation and correction of the needle path. A typical brachytherapy setup, diagrammed in Fig. 1, includes an ultrasound probe registered to a guide template that consists of a grid of holes. Each brachytherapy needle is inserted through a grid hole that corresponds to a pre-planned target location in the prostate. During needle insertion, an ultrasound image at the specified target depth is acquired to check the needle position. Only a small portion of the needle can be imaged at any time as the field of view of the ultrasound is very narrow . Thus, as the needle is sent off course from its target, the surgeon is unable to see or predict the needle\u2019s total deflection", + " An 18-gauge by 200 mm prostate seeding needle (Eckert & Zielger BEBIG GmbH, Berlin, Germany) is inserted through a brachytherapy template (Model D0240018BK, C. R. Bard, Inc., Covington, USA) into a tissue phantom. The tissue phantom is a plastisol gel from M-F Manufacturing Co, Fort Worth, USA. The physical characteristics of the plastisol gel mimic human tissue with the added benefit that the phantom tissue transmits both visible light and ultrasound waves, allowing simultaneous capturing of images in both modalities. A portion of the needle is imaged by a 4DL14-5/38 Linear 4D transducer (different from the transrectal ultrasound probe shown in Fig. 1) which is connected to an SonixTouch ultrasound machine (Ultrasonix, Richmond, Canada). For this experiment, we only use the 2D functionality of the ultrasound probe. The entire needle is imaged from above by a XCD-SX90CR video camera (Sony Corporation of America, New York, USA). The prostate seeding needle is inserted such that it deflects in the plane imaged simultaneously by the camera and the ultrasound probe. Due to the process by which ultrasound and camera images are captured, the ultrasound probe does not interfere with the camera\u2019s field of view even while they both image the same plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002840_10402004.2020.1717703-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002840_10402004.2020.1717703-Figure4-1.png", + "caption": "Figure 4. Schematic side views of disks showing (a) the nozzle configurations of the oil delivery system and (b) temperature measurement systems.", + "texts": [ + " The torque limiting couplings and flex couplings \u2465 on the motor side of the torque meter were able to isolate any adverse effects of typical misalignments on torque measurements. A high-temperature and high-pressure lubrication system was implemented in the machine of Figure 3(a). Two electric sump heaters were used to heat the lubricant in order to provide elevated lubricant temperatures. A pump circulated the lubricant through the entire lubrication Acc ep te d M an us cr ipt system to deliver lubricant at inlet temperatures up to 160\u00b0C. The inlet manifold shown in Figure 4(a) housed the thermocouple used to control inlet lubricant temperature and two lubricant outlet ports. The outlet ports could either be plugged or contain an oil jet. The torques provided to both disks via their spindles were measured by the inline torque meters as signals 1T and 2T . They are different from the traction torques ,1fT and ,2fT shown in the free-body diagrams of Figure 1, which counter the moment caused by surface traction. iT are the torques applied to the entire shaft assemblies to keep them in equilibrium", + " In typical arrangements, the probe is spring loaded against the measurement surface to assure continuous surface contact, in the process potentially generating additional frictional heat at the sensor-disk interface to skew the measurement. Likewise maintaining a continuous contact between the sensor and the disk surface without increasing the spring load is a concern. In an effort to reduce frictional heat generation, a thermocouple fixture was designed to hold the thermocouple in place and center it in the contact area of the rotating disk. A rigid surface thermocouple was selected so that the device would stay in contact with the surface. As shown in Figure 4(b) schematically, the mounting fixture was placed such that the thermocouple was slightly loaded against the disk at an angular position of 225\u00b0 out of contact. Given the critical importance of the measurement of the surface bulk temperatures and potential liabilities of the contact-type thermocouples, another redundant measurement system was implemented. This system employed a non-contact temperature sensor, placed 270\u00b0 out of contact as shown in Figure 4(b). Emitted radiation and emissivity measurements of the sensor are related to corresponding surface temperatures through calibration. As this method relies on clear sight of the surface, presence of oil and mist formed at high rotational speeds is likely to influence its measurements. To ease some of these issues, a plastic wiper was mounted close to Acc ep te d M an us cr ipt the disk surface to whisk away any excess oil, and a low-pressure air purge was added to clear the measurement surface from oil as much as possible", + " The surface absorbs and reflects radiation, and \u03b5 must be known in order to convert the acquired signal to surface temperature accurately. Typical values used for polished steel surfaces were found to be 0.12 . The disk temperature was elevated to various set values in a specimenheating oven and the non-contact sensor measurements were compared to those from a calibrated contact thermocouple to establish the proper \u03b5 value. The measurement range of the infrared thermometer used in this study was 100-600 \u00b0C. Readings of this sensor at temperatures below 100 \u00b0C are not accurate. It is noted In Figure 4(b) that only one of the disks (flat disk 2) was instrumented to measure ,2b in this set-up in order to minimize instrumentation on the radially moving disk 1 spindle. Bulk temperatures of disks, ,1b and ,2b , cannot be assumed to be equal since 1 2 in general (11,22). A method was established to estimate ,1b with the instrumentation described above. As both disk-spindle assemblies are identical (i.e. both spindles have the same geometries with identical bearings carrying the same amount of radial load, and disk blanks have the same Acc ep te d M an us cr ipt geometries with the only difference being the axial crown of disk 1) convective cooling of both spindles was assumed to be the same at equivalent i " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002879_s10878-020-00533-z-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002879_s10878-020-00533-z-Figure7-1.png", + "caption": "Fig. 7 Tokens and robots movement. a Initial state of the SCS. At time 0, robot ui holds token zi . b State of the system at time t1 > 0. Robots u0 and u3 met and exchanged tokens. c State of the system at time t2 > t1 > 0. Robot u3 and its token z0 have been removed", + "texts": [ + " In this section we introduce the notion of a token as an abstract entity used to describe the behavior of a partial SCS. The key idea is to focus on the tokens instead of the robots. This will be crucial to compute the broadcasting resilience. Let u0, . . . , un\u22121 be the set of robots at the beginning of time (when the SCS is deployed) each one carrying a token. We assume the following protocol: at all times, each active robot holds a token. When two robots meet, they exchange their tokens. When a robot is removed, the token it carries is removed as well, see Fig. 7. Remark 5 A starving robot never exchanges its token. Lemma 2 Consider a partial SCS. Independently of the number of removed robots and when they were removed, a live token always remains in its initial ring and moves with constant speed (2\u03c0 per unit time) along its ring. Proof Let z be a token in a partial SCS. Let u be the robot bearing z at an arbitrary time t . Let r be u\u2019s ring at time t . Let t > 0 be a real number. Suppose first that in the time interval from t to t + t , robot u remains active and does not reach any link position", + " Remark 6 The previous lemma states that the relation of correspondence between two tokens remains invariant unless one of them is removed from the system. Definition 11 (Token graph) Let F be an m-partial SCS. Let Z be the set of surviving tokens. The token graph of F is the graph TF whose vertices are the tokens in Z [that is, V (TF ) = Z ] and whose set of edges is: E(TF ) = {{z, z\u2032}|z \u2208 Z , z\u2032 \u2208 Z , there is a correspondence between z and z\u2032}. From Remark 6, the token graph of a partial SCS remains invariant while no additional robots removed. Figure 9a shows the token graph of the SCS of Fig. 7 before the robot u3 is removed. From the definition of token graph and Lemma 5, the next result (illustrated in Fig. 9b) follows. Lemma 6 Let F be an m-partial SCS with set of tokens Z. Let F \u2032 be the m\u2032-partial SCS resulting from the removal of some robots from F , with Z \u2032 \u2282 Z, the set of tokens of F \u2032. Let TF and TF \u2032 be the token graphs of F and F \u2032 respectively. Then, TF \u2032 is the subgraph induced by Z \u2032 in TF . The following lemma relates starvation with token and will be useful for computing the k-isolation resilience" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002399_jcej.15we235-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002399_jcej.15we235-Figure1-1.png", + "caption": "Fig. 1 The rheological experiment measuring system Fig. 2 Flow condition of lubricating grease", + "texts": [ + " Pan (E-mail address: jbpancumt@163.com). Research Paper 816 Journal of Chemical Engineering of Japan peratures are obtained. The research results could provide a fundamental basis for development of new centralized lubrication system. Sample of NLGI 3 lithium grease, which was selected as the representative sample in this study, was commercial products manufactured by Lubricant Tianjin Company, Sinopec Lubricating Oil Co., Ltd. (China). Its main components and technical data are listed in Table 1. Figure 1 displays the schematic diagram of parallel-plate measuring system (Anton Paar GmbH, 2011). It was applied to measure the viscoelastic properties and flow properties of NLGI 3 lithium grease by using a plate\u2013plate geometry (50 mm diameter, 1 mm gap) (Delgado et al. 2005a, 2006; S\u00e1nchez et al., 2011). The peltier heating system was used in the rotational rheometer. In this work, due to the potential thermal degradation at high temperatures, the temperature ranging from 25 to 85\u00b0C was selected as the measuring temperature (Cann et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003961_j.measurement.2020.108909-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003961_j.measurement.2020.108909-Figure1-1.png", + "caption": "Fig. 1. Construction of IMD DBB.", + "texts": [ + " Commonly, the circular test underestimates the errors in comlex geometry and the measuring values thereof are less reliable to be used as a quantitative indicator for the accuracy of a manufacturing or manipulating process with 3 axes or more involved. In order to develop an efficient measuring approach to acquire more reliable information about machine accuracy under shop floor conditions, we devoted our efforts in our former work to the following two aspects: \u2022 we constructed a Ballbar with enhanced measuring range (see Fig. 1) and \u2022 we developed a universal approach to algorithmically generate spatial measuring paths,3 which are feasible, smooth, free of collision and at the same time sensitive to the kinematics of the machine concerned. For the sake of completeness, we will first briefly present the measuring instrument and the path generation algorithm in the next section. Then we will address the necessity of determining the correspondence between the actual poses of the machine concerned and the Ballbar measuring values for quantitative tasks", + " Continuous spatial measurement with Ballbar The commercially available telescoping Ballbar normally has very limited measuring range (e.g. \u00b11 mm by Renishaw QC20-W [10]), which does not suffice for the continuous spatial DBB measurement according to our experience. Hence, a Double Ballbar with enhanced measuring range was designed and manufactured at Institute of Mechatronic Engineering, TU Dresden (IMD). Unlike the common commercial Double Ballbars with transducer, IMD DBB uses an optical measuring principle. As Fig. 1 shows, the optical sensor head scans the scale tape and builds an incremental measuring system with a resolution of 0.1 \u03bcm. We experimentally validated in our work [23] that the IMD DBB could reach a comparable accuracy (1 \u03bcm in 5 mm measuring range) to the commercial DBB (\u2264 1 \u03bcm in \u00b10.1 mm measuring range by QC20W [10]). Fig. 2 visualizes the physical object of IMD DBB, the material of which is under 1000 euro cost range. Before each measurement, the DBB would be calibrated with an external calibrator (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003320_j.cmpb.2020.105646-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003320_j.cmpb.2020.105646-Figure11-1.png", + "caption": "Fig. 11. The movement of the digital articulator (a) Opening (b) Protrusion (c) Lateral excursion.", + "texts": [ + " Secondly, we move (open) the mechanical articulator and fix its position with thermoplastic material, and use the scanner to scan the upper teeth and lower teeth. The position information of two casts can provide the transformation from the closed position to the moved position. Meanwhile, we track the markers which are attached to the mechanical articulator to measure the position of the incisal pin tip which can provide the corresponding transformation by the kinematic model we analyze. Comparing the two transformations, we obtain the error between the digital and mechanical articulator. We use five poses for the measurement ( Fig. 11 ). As the verification method we previously described, we can measure the transformation matrix from initial to final pose of the real and digital articulator. After decomposing the transformation matrix into 6 different parameters and comparing them, the result is shown in Table 1 . In the next phase of our research we will use the same optical tracking system to capture the dynamic occlusion of a real patient and the multi-face markers need to be placed outside of the patient\u2019s face and mouth, one close to the upper jaw, and the other close to the lower jaw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-FigureA.37-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-FigureA.37-1.png", + "caption": "Fig. A.37. Main parameters of the Waveboard multibody model and body frames.", + "texts": [ + " (55) , being the numerical values of the trajectory parameters provided in Table 2 : y 2 ( t ) = A 1 sin ( \u03c9t ) , \u03b12 ( t ) = A 2 sin ( \u03c9t ) , \u03b32 ( t ) = A 3 sin ( \u03c9t ) , \u03b843 ( t ) = t, \u03b852 ( t ) \u03b15 ( t ) \u2212 \u03b12 ( t ) = A 4 sin ( \u03c9t \u2212 \u03b4) \u2212 A 2 sin ( \u03c9t ) . (55) In Eq. (55) , the relative angle \u03b852 between the rear and front decks is computed as the difference between \u03b15 and \u03b12 , which represent approximately the twist of the bodies 5 and 2, respectively. The parameter \u03b4 has been used to consider the existing gap between both platforms. The numerical simulation is performed considering a value of the casters angle of inclination \u03b8 c and torsional stiffness k t (see Fig. A.37 in Appendix A) of \u03b8c = 30 \u25e6 and k t = 100 N m rad . As in the case of the direct dynamics simulation, a numerical value of \u03b1e = 50 has been used for the Baumgarte stabilization constant. Fig. 13 shows the time evolution of forces F 2 x , F 2 y and F 5 y . Indeed, F 2 y and F 5 y have an oscillatory variation with time, being the amplitude of the forward lateral action larger. Moreover, F 2 x is significantly lower than F 2 y and F 5 y , though it has a nonzero value. In future works the control of the underactuated Waveboard with four actuations ( F 2 y , F 5 y , M 2 x and M 5 x ), avoiding the use of F 2 x , will be addressed as an optimal control problem to look for controls that drive the system by numerical optimization", + " Garc\u00eda-Vallejo: Conceptualization, Methodology, Formal analysis, Investigation, Writing - original draft, Writing - review & editing. E. Freire: Conceptualization, Methodology, Formal analysis, Investigation, Writing - original draft, Writing - review & editing. This work was supported by Grant FPU18/05598 of the Spanish Ministry of Science, Innovation and Universities. The authors also acknowledge the financial support received from foundation Manuel Gayan Buiza. The geometric parameters and body frames of the model are shown in Fig. A.37 . Starting from the center of mass of body 2, the absolute position vectors of the centers of mass of the rest of the bodies (origins of the body frames) can be obtained as follows: r G 2 = ( x 2 y 2 z 2 )T , r G 3 = r G 2 + R 21 \u0304r 3 , r G 4 = r G 3 + R 31 \u0304r 4 , r G 5 = r G 2 + R 21 \u0304r 5 , r G 6 = r G 5 + R 51 \u0304r 6 , r G 7 = r G 6 + R 61 \u0304r 7 , (A.1) where the orientation matrices of the body frames are given by: R 21 ( q ) = R \u03b12 R \u03b22 R \u03b32 , R 31 ( q ) = R 21 R int R \u03b832 , R 41 ( q ) = R 31 R \u03b843 , R 51 ( q ) = R 21 R \u03b852 , R 61 ( q ) = R 51 R \u03b865 , R 71 ( q ) = R 61 R \u03b876 , (A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001049_ijat.2013.053164-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001049_ijat.2013.053164-Figure9-1.png", + "caption": "Figure 9 Steady-state chip formation and tool wear development (see online version for colours)", + "texts": [ + " When the critical drill geometries are accurately regenerated, a constant continuation of the preceding drilling path is maintained in the successive cycle of drilling. The process is stable with the development of steady drill degradation and a straight hole should be produced eventually. Under such condition, tool wear is developed steadily on different parts of the carbide drill tip throughout the drilling cycle. Wear regions include the rake and flank faces of the cutting edges, both inner and outer; the bearing pads and side margin which is part of the bearing configuration. As depicted in Figure 9, chip formation commences as the rotating drill tip engages with the workpiece at the hole bottom created in the preceding cycle. The drill apex touches the preceding hole profile at early stages and then harnessed the drill tip into the workpiece. Chip is thus firstly formed around the drill apex. When the drill is fed into the workpiece further, material removal intensifies with the full engagement of the inner and outer cutting edges. This is accompanied with chip growth and force generation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001900_imece2013-63817-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001900_imece2013-63817-Figure2-1.png", + "caption": "FIGURE 2. SAMPLE OF PRODUCTS MADE IN PREVIOUS YEAR. TOP: (LEFT) STAIR CLIMBING DOLLY (RIGHT) TENNIS BALL LAUNCHER. BOTTOM: (LEFT) PART OF A COMPRESSOR (RIGHT) PARTS OF A SWISS ARMY TYPE KNIFE", + "texts": [ + " Propose automation strategy for large volume manufacturing. In the beginnings of the semester, the entire class participates in a product selection process. Selection is based on criterion such as the market potential for the product, complexity of the design, possibility for innovation, number of components in the product etc. The products developed by the class in the past years include a combination lock, a multi- purpose tool, turbo-charger, power tool, stair climbing dolly, tennis ball launcher to name a few (Figure 2). Even though each student is a member of only one group, they are responsible for the over completion of the product realization process and hence are required to interact closely with other groups to make this happen. Figure 1 shows how groups interact among themselves. Interactions happen within class, outside the class and through on-line medium. Every week, each group gives a brief update of their progress. Groups utilize this opportunity to exchange information. Figure 3 shows the time line and the work load for each group" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001388_isie.2014.6864934-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001388_isie.2014.6864934-Figure5-1.png", + "caption": "Fig. 5. Typical installation of the smart sensor, embedded a HV disconnect switch.", + "texts": [ + " The discharging process was tested under different operation modes, with the following results: \u2022 In case the sensor was programmed to communicate every 10s, the energy stored in the fully charged capacitors would allow operation during 2 hours and 30 minutes; \u2022 In case it was programmed to communicate every hour, operation would last for 10 hours; \u2022 In the standard factory programming of this smart sensor, temperature is hourly monitored, and communication occurs twice a day in normal circumstances, which corresponds to power autonomy of 15 hours. Fig. 4 shows three discharge curves, one for each operation mode. B. Optoelectronic Sensor The alignment between the two parts of a disconnect switch is measured by the optoelectronic sensor located in one part only (female), as illustrated in Fig. 5, the optical coupling being ensured by a reflective surface located in the opposite male part, as both parts rotate in the same plan Fig. 6. Thus, whenever both arms in a HV switch sensing device and the reflective surface are each other (alignment angle 0\u00ba) and only phototransistors are irradiated with infra-red light two parts become misaligned, the angle between them implies that the reflected infra-red beam irradiates other than the central ones, thus allowing the degree of misalignment to be determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001615_j.isatra.2015.05.018-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001615_j.isatra.2015.05.018-Figure2-1.png", + "caption": "Fig. 2. The input commands regarding the inertial reference frame involves two direction angles entitled as \u03b1 and \u03b2.", + "texts": [ + " (i) It is applicable on the mode of engine off to deal with the pointing process and (ii) it is also applicable on the mode of engine on to deal with the finite burn orbital transferring. In order to present the investigations, in a better form, the schematic of the orbital transferring from initial orbit to its final orbit based upon intersection radius, in finite burn, is easily illustrated in Fig. 1. Now, to present the control strategy in engine on mode, the input commands have to be given as illustrated in Fig. 2. In one such case, T- is given as the thrust vector of spacecraft, which is relative to the inertial coordinate system. In the same way, the conversion of thrust from vector to its scalar representation regarding the inertial reference frame involves two direction angles entitled \u03b1 and \u03b2; respectively. In fact, the parameters \u03b1 and \u03b2 are generated through a guidance system to guarantee the process of transferring from the initial orbit to its final one. Moreover, the present results can easily be converted to Euler angles via Phicom \u00bc 0 Thetacom \u00bc \u03b1 Psicom \u00bc \u03b2 8>< >: \u00f01\u00de The schematic diagram of the proposed SPCMAC strategy is now illustrated in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure8.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure8.6-1.png", + "caption": "Fig. 8.6 Left: 3D sketch for the contact forces: vertical restrictive force, lateral friction and the power torque in the first wheel. Right: Sketch of the position of local reference frames at the contact point of each wheel of a differential drive robot, using regular wheels", + "texts": [ + " . . . . . . . . . 337 Fig. 8.4 Sketch of the virtual reference frames position at the contact points of each of the three omni-wheels of an omnidirectional robot . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 337 Fig. 8.5 Differential robot Pioneer 3DX from Adept Technology, Inc. for in-door purposes, [Photograph taken from http://www.mobilerobots.com/Libraries/Downloads/ Pioneer3DX-P3DX-RevA.sflb.ashx] . . . . . . . . . . . . . . . . . . . . . . 346 Fig. 8.6 Left: 3D sketch for the contact forces: vertical restrictive force, lateral friction and the power torque in the first wheel. Right: Sketch of the position of local reference frames at the contact point of each wheel of a differential drive robot, using regular wheels . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 347 Fig. D.1 Parallelepiped with Cartesian coordinates. x-axis: \u00bd a; a with a total length of A \u00bc 2a; y-axis: \u00bd b; b with a total length of B \u00bc 2b; and z-axis: \u00bd c; c with a total length of C \u00bc 2c ", + "5, which has two regular driving wheels and a castor wheel to provide stability in the floor plane. For kinematic analysis, consider the local non-inertial frame to be placed at the geometric center of the two driving wheels such that the x-axis is positive to the front of the vehicle, the z-axis is positive in the upward direction and the y-axis is placed according to the right-hand rule. For simplification purposes, consider the wheels to be mass-less bodies and the contact point between the floor and the wheel to be constant relative to the wheel position (refer to Fig. 8.6-Left). The lateral friction of the castor wheel is neglected. Consider a local frame ri at the contact point of each driving wheel i with the x-axis parallel to the rotation of the wheel and properly oriented such that a positive torque \u03c4mi provides a positive displacement of the robot. For the castor wheel consider a frame r3 at the constant point of its contact 3 that is parallel to the non-inertial frame of the robot, i.e. Rr3 1 = I , while the Rotation matrices or the driving wheel contact frames are just a simple rotation of \u03c0/2 rad around the vertical (z-axis), i", + " These operators takes the following shape (next page): X T 1 = T T (rr1)R r1 1 T X T 2 = T T (rr2 )R r2 1 T = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 0 1 0 0 0 0 \u22121 0 0 0 0 0 0 0 1 0 0 0 \u2212h 0 d 0 1 0 0 \u2212h 0 \u22121 0 0 0 \u2212d 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 ; = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 0 1 0 0 0 0 \u22121 0 0 0 0 0 0 0 1 0 0 0 \u2212h 0 \u2212d 0 1 0 0 \u2212h 0 \u22121 0 0 0 d 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 ; 348 8 Model Reduction Under Motion Constraint X T 3 = T T (rr3)Rr3 1 T = \u23a1 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 1 0 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 0 0 h 0 1 0 0 \u2212h 0 L 0 1 0 0 \u2212L 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 From Fig. 8.6-left and the fact that only the castor wheel is omnidirectional, the 3 wrenches expressed in local coordinates ri are given as: F(r1) c1 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d fcx fcy fcz \u03c4cx \u03c4cy \u03c4cz \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d fx1 1 r \u03c4m1 fz1 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 ; F(r2) c2 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d fx2 1 r \u03c4m2 fz2 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 ; F(r3) c3 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 fz3 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 (8.45) where r is the radius of the driving wheels, and the torques \u03c4mi are those produced by the motors at them. In this case, each wheel including the castor one produces vertical constraint forces fzi that constraint the robot to remain in the floor level; but only the driving wheels produce a lateral friction forces that constraint the robot from lateral slippering" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003261_tii.2020.3004389-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003261_tii.2020.3004389-Figure4-1.png", + "caption": "Fig. 4 The arrangement of round conductors in the slot windings.", + "texts": [ + " The winding Authorized licensed use limited to: University of Exeter. Downloaded on June 26,2020 at 10:25:00 UTC from IEEE Xplore. Restrictions apply. 1551-3203 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. IEEE TRANSACTIONS ON INDUSTRIAL INFORMATICS 3 3 loss is separated into the slot and the end parts in the following analysis because of the difference in flux amplification effect. Fig. 4 presents the cross-section of the slot winding and a possible arrangement of conductors. To facilitate the analysis below, the wires in the slots are divided into 15 layers along the radial direction. Assuming the induced current is resistance limited and the magnetic field generated by eddy current is negligible compared to the external field, the additional eddy losses caused by the stator field in each conductor of the l th winding layer can be estimated by [20] 4 2 0 21 , ( ) 128 a l Eddy act l NId L P w \u2212= (1) in which, is the copper electrical conductivity, 0 is the space permeability, is the electrical angular speed of the input current, d is the radius of the bare wire, La is the length of the active part, l w is the slot width at the l-th layer, and 1l NI \u2212 is the sum of current in the conductors below the l-th layer", + " The estimation result proved to be sufficiently accurate where the conductor dimension is smaller than the skin depth. The ac loss is usually expressed in terms of a dimensionless ac loss factor Kac to represent the growth on the dc loss: , , , ,ac ac act dc act ac act dc act K P P R R= = (2) where ,ac act R and ,dc act R are the equivalent ac and dc resistances of the slot winding. Their values are estimated according to the length ratio of conductor in the slot and end winding parts. For the l-th winding layer shown in Fig. 4, a lateral fill factor l is defined as the copper filling ratio. 2 4 l l l N d w d = (3) in which, l N is the number of conductors in the l-th layer. For each conductor in the l-th winding layer, the ac loss factor ac K can be expressed as 2 4 1 0 1 1 4 l ac l l NI d K N I \u2212 = + (4) in which, 0 I is the amplitude of input current, is the skin depth at the electrical angular speed 0 2 = (5) The winding losses caused by the rotation of the rotor are calculated using full-scale FE models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure5.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure5.3-1.png", + "caption": "Fig. 5.3 Complex body composed of basis shapes", + "texts": [ + " It is important to note that this center of mass does not depend on the inertial frame because it does not depend on the attitude of the body. Even more, for rigid bodies, is a constant fixed point relative to the body\u2019s frame 1. Then both the constant mass m and the constant position of the center of mass rc are lumped parameters of the rigid body B. For examples refer to Appendix D Remark It is easy to see that for bodies of homogenous density, each coordinate of the center of mass vector is null if the body is symmetric along the corresponding axis. Example 5.1 (Complex body) Consider the multibody system of Fig. 5.3, with a common reference frame at the centroid of the ellipsoid; with the following values: 5.1 The Center of Mass 237 Ellipsoid Pole Cylinder Box m1 = 4 kg m2 = 0.2 kg m3 = 0.11 kg m4 = 4 kg A1 = 0.6m A2 = 0.7m r3 = 0.012m A4 = 0.45m B1 = 0.2m r2 = 0.0085m B3 = 0.05m B4 = 0.25m C1 = 0.3m C4 = 0.3m Notice that the center of mass vectors for each of the four bodies, with coordinates of the local frame placed at the centroid of the ellipsoid are: rc1 = \u239b \u239d00 0 \u239e \u23a0 ; rc2 = \u239b \u239d\u2212A1 2 \u2212 A2 2 0 0 \u239e \u23a0 = \u239b \u239d\u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000723_s10846-013-0002-9-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000723_s10846-013-0002-9-Figure2-1.png", + "caption": "Fig. 2 Kinematic model of unicycles", + "texts": [ + " The second one is devoted to propose two different solutions to the n\u2212trailer emulation problem in continuous-time. The third one contains two different solutions based on the first order discrete-time approximation of the kinematic model of the agents. Section 4 includes the description of the experimental platform as well as numerical and real time experiments. Finally, some conclusions and outlooks are offered in Section 5. 2.1 Kinematic Model of Unicycles Denote by {R1, . . . , Rn} a set of n differentially driven mobile robots moving in the plane. The kinematic model of each robot Ri , as shown in Fig. 2, is given by \u23a1 \u23a3 x\u0307i y\u0307i \u03b8\u0307i \u23a4 \u23a6 = \u23a1 \u23a3 cos \u03b8i 0 sin \u03b8i 0 0 1 \u23a4 \u23a6 [ vi wi ] , i = 1, . . . , n, (1) where (xi, yi) are the coordinates of the mid-point of the wheels axis, \u03b8i is the orientation of the robot with respect to the X-axis, vi is the longitudinal velocity of the mid-point and wi is the angular velocity. It is known [5] that the dynamic system (1) can not be stabilized by any continuous and time-invariant control law. To overcome this obstruction in the literature it is common to study the kinematics of a point \u03b1i = (pi, qi) T outside the wheels axis [17, 23]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001071_s11740-013-0466-2-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001071_s11740-013-0466-2-Figure3-1.png", + "caption": "Fig. 3 Kinematic modulation in cylindrical grinding", + "texts": [ + " In addition, grinding marks vertical to the rolling direction can be manufactured by using a side grinding process, so that comparable contact conditions of gear pairs can be simulated in the two-disk test rig. The influence of kinematic modulation in abrasive machining was systematically investigated for the analogous process. The experimental tests were carried out using an universal cylindrical grinding machine by Schaudt Mikrosa BWF GmbH, Stuttgart, Germany. Due to good damping properties of the machine, dynamic kinematic modulations were particularly well realized. A schematic sketch of the kinematic modulated cylindrical grinding process of bearing rings is shown in Fig. 3. In the experimental tests a cooling lubricant emulsion was used. Whereas grinding feed rates vfa up to 15,000 mm/min were achieved with the high-precision machine, frequencies fs up to 2.3 Hz were used for the cylindrical grinding process of bearing rings. In the experimental tests the frequency fs and the amplitude As were varied for the kinematic modulated cylindrical grinding process, displayed in Table 1. Further setting parameters like grinding wheel circumferential speed vs and speed ratio q as well as the dressing parameters for the vitrified bonded CBN grinding wheel were not varied during the grinding tests", + " Figure 8 presents the simulation results of the kinematic modulated grinding by varying the frequency fs relating to the surface parameters Sa, Sq, Sz as well as the ratios Spk/Svk and Sk/Sz. The influence of a kinematic modulation is clearly recognizable. An increase of the frequency up to fs = 2.3 Hz leads to a reduction of the surface parameters. A rising frequency fs without changing the feed rate vfr is resulting in a higher stroke velocity and therefore the cutting speed vc is increasing, Fig. 3. With respect to the ratio Spk/Svk from the Abbott curve, the surface structure shows appropriate properties for rolling contacts by using a frequency fs = 2 Hz. By contrast, no significant differences in the ratio Sk/Sz are recognizable for different frequencies. In another experimental test, the influence of the variation of the amplitude As in kinematic modulated grinding was investigated. A kinematic modulation results in better surface qualities of the workpiece in the simulation as well as in the grinding test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003697_ecce44975.2020.9235609-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003697_ecce44975.2020.9235609-Figure13-1.png", + "caption": "Fig. 13. Total deformation of the frame surface: (a) Conventional; (b) W-OCC", + "texts": [ + " The proposed current profiles obtained from the optimum multi-objective control algorithm are applied to 3D Electromagnetic FEA, and the harmonic forces are calculated on stator teeth. The harmonic forces obtained from electromagnetic analysis is imported to harmonic response analysis. In the harmonic response analysis, radial deformation, acceleration, and surface velocity are calculated on the outer surface of housing in the frequency domain. The maximum deformation of the frame surface at around 2900 Hz is 0.05 \u00b5m for the conventional current profile and 0.037 \u00b5m for the proposed W-OCC current profile as presented in Fig. 13. (a) (b) Fig. 14. Acceleration of frame surface: (a) Conventional; (b) W-OCC The acceleration levels are obtained from mechanical analysis as shown in Fig. 14. The maximum radial acceleration of the frame surface at 2900 Hz is 1.4 m/s2 for the conventional current profile and 0.85 m/s2 for the proposed WOCC current profile. VII. EXPERIMENTAL SETUP AND TEST RESULTS The test bench has prototype SRM in Fig. 15, 1000 V battery simulator, Horiba Dyno systems, and SRM inverter. The control method is implemented using the TILAUNCHXL-F28379D DSP controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000629_ieeegcc.2013.6705835-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000629_ieeegcc.2013.6705835-Figure1-1.png", + "caption": "Fig. 1: Body- xed and inertial reference frames.", + "texts": [ + " Finally, we formulate the SFDC problem for the linearized model of the AUV. This section describes the kinematic and dynamic equations of motion of an AUV. The dynamic model of an AUV is developed through Newton-Euler formulation by using conservation laws of linear and angular momentum. The equations of motion of such vehicles are highly nonlinear [17] and coupled due to the hydrodynamic forces which act on the vehicle. The equations of motion of an underwater vehicle in six degrees of freedom with respect to the body- xed frame (as shown in Fig. 1) can be written as [17]: M\u03bd\u0307 + C(\u03bd)\u03bd +D(\u03bd)\u03bd + g(\u03b7) = \u03c4, (1) where \u03bd = [u, v, w, p, q, r]T denotes the vector of linear and angular velocities in the vehicle coordinate frame and \u03b7 = [x, y, z, \u03c6, \u03b8, \u03c8]T denotes the vector of absolute positions and orientations (Euler angles) in the inertial frame, M = MRB + MA and C(\u03bd) = CRB(\u03bd) + CA(\u03bd) are the system inertia matrix (including added mass) and the Corioliscentripetal matrix (including added mass), respectively. D(\u03bd) denotes the resulting matrix of linear and quadratic drag, g(\u03b7) denotes the vector of gravitational/buoyancy forces and moments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001873_0954406215585186-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001873_0954406215585186-Figure16-1.png", + "caption": "Figure 16. The bearing test rig.", + "texts": [ + " The PE values and the SNR values quantitatively indicate that the transient signal extracted by the proposed method has a better data quality. Case 2: A bearing inner race fault signal obtained from an experimental motor The experimental setup of the bearing defect test. A bearing signal with slight inner race fault was used to further validate the proposed method. The bearing data used in this section were obtained from the Case Western Reserve University Bearing Data Centre Website.25 The dataset was acquired by using the test rig as shown in Figure 16. The test rig consists of a torque transducer, a 2 hp motor, a dynamometer, and the control electronics (not shown). The test bearings were used to support the motor shaft. The vibration signals were collected using accelerometers mounted on the housing with magnetic bases. The sampling frequency was set 12 kHz for fan end bearing experiments. The deep groove ball bearings 6205-2RS JEM SKF were used in this test. The details of the structural parameters of this type of bearing are given in Table 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure6.13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure6.13-1.png", + "caption": "Figure 6.13 A curve-on-curve joint.", + "texts": [], + "surrounding_texts": [ + "(a) Relative Position Equations The fact that Qab and Qba are coincident points can be expressed as follows: r\u20d7abq = r\u20d7ab,ba + r\u20d7baq (6.124) The vectors in Eq. (6.124) can be represented by the following column matrices in ab and ba. r(ab) ab,ba = uixab + ujyab + ukzab (6.125) 148 Kinematics of General Spatial Mechanical Systems r(ab) abq = uifabi(pa) + ujfabj(pa) + ukfabk(pa) (6.126) r(ba) baq = uifbai(pb) + ujfbaj(pb) + ukfbak(pb) (6.127) Hence, Eq. (6.124) can be replaced with the following matrix equation. ui[fabi(pa) \u2212 xab] + uj[fabj(pa) \u2212 yab] + uk[fabk(pa) \u2212 zab] = C\u0302(ab,ba)[uifbai(pb) + ujfbaj(pb) + ukfbak(pb)] (6.128) In Eq. (6.128), the orientation matrix can be expressed as follows with suitably selected constant unit column matrices nab1, nab2, and nab3: C\u0302(ab,ba) = en\u0303ab1\ud835\udf19ab en\u0303ab2\ud835\udf03ab en\u0303ab3\ud835\udf13ab (6.129) As seen above, the curve-on-curve joint involves eight variables, which are xab, yab, zab;\ud835\udf19ab, \ud835\udf03ab, \ud835\udf13ab; pa, pb The eight variables listed above are interrelated by three independent scalar equations contained in Eq. (6.128). Therefore, the mobility of this joint is \ud835\udf07ab = 8\u2212 3 = 5. Thus, if five of the variables (e.g.\ud835\udf19ab, \ud835\udf03ab,\ud835\udf13ab, xab, yab) are specified, then the other three variables (i.e. pa, pb, and zab) can be determined by using Eq. (6.128). (b) Relative Velocity Equations The differentiation of Eq. (6.128) leads to the following relative velocity equation. ui(f \u2032abip\u0307a \u2212 x\u0307ab) + uj(f \u2032abjp\u0307a \u2212 y\u0307ab) + uk(f \u2032abkp\u0307a \u2212 z\u0307ab) = ?\u0303? (ab) ab,ba[ui(fabi \u2212 xab) + uj(fabj \u2212 yab) + uk(fabk \u2212 zab)] + p\u0307bC\u0302(ab,ba)(uif \u2032bai + ujf \u2032baj + ukf \u2032bak) (6.130) In Eq. (6.130), \ud835\udf14(ab) ab,ba is obtained from Eq. (6.129). That is, \ud835\udf14 (ab) ab,ba = ?\u0307?abnab1 + ?\u0307?aben\u0303ab1\ud835\udf19ab nab2 + ?\u0307?aben\u0303ab1\ud835\udf19ab en\u0303ab2\ud835\udf03ab nab3 (6.131) Equation (6.130) can be used so that, if five of the derivatives (e.g. ?\u0307?ab, ?\u0307?ab, ?\u0307?ab, x\u0307ab, y\u0307ab) are specified, then the other three derivatives (i.e. p\u0307a, p\u0307b, and z\u0307ab) can be determined. However, it is to be noted that, although the curve-on-curve joint is a rolling contact joint, it is not possible to have a relative rolling-without-slipping motion between its sharp-edge kinematic elements. If there is sticking friction between these kinematic elements, then the relative motion becomes similar to that between the kinematic elements of a spherical joint." + ] + }, + { + "image_filename": "designv11_34_0003694_ecce44975.2020.9235404-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003694_ecce44975.2020.9235404-Figure17-1.png", + "caption": "Fig. 17. Flux density and magnet loss distributions in the skewed stator of the integrated motor-compressor.", + "texts": [ + " In the unaligned stator-rotor configurations with a 38-degree rotor stagger angle, with the stator-rotor skew angle increases from 0 degrees to 38 degrees, the average loaded torque decreases by 25% due to the unalignment between the stator and rotor, which induces more leakage flux paths and reduces the flux linkage. The cogging torque decreases with the increase of the stator-rotor skew angle as well. d. Magnet and iron losses Iron loss and magnet losses have a significant effect on motor efficiency, especially for high-speed motors with high fundamental frequency [19],[20]. In the integrated motorcompressor, the rotor, stator, and magnets are skewed. As the reluctance of the magnet is much larger than the laminations, in a skewed magnet shown in Fig. 17, the flux flows in a minimum reluctance path across the magnet, which is perpendicular to the edges of the skewed magnets. In a skewed stator core, the flux flows close to the left edge of the stator core in Fig. 17 as the flux tries to flow across the magnet in the direction perpendicular to the boundaries between the magnet and stator core. When the flux crosses the magnet at the right corners of the boundary between the magnet and stator core, the direction of the flux is also perpendicular to the boundary when the flux enters the magnet. However, in that case, the flux path across the magnet is longer than the path of the flux entering the magnet from the left boundary due to the skewed shape of the magnets", + " Cogging torque and average rated torque with respect to the stator-rotor skew angle in the unaligned stator-rotor configurations with stator-rotor skew angle from 0-38 degrees. 0 10 20 30 40 Stator-rotor skew angle [deg] -0.5 0 0.5 1 5550 Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 22:42:22 UTC from IEEE Xplore. Restrictions apply. Therefore, the flux in the skewed stator cores tends to avoid the right corners of the magnets, which leads to unevenly distributed flux density in the stator cores and magnets, and hence the higher magnet eddy current losses shown in Fig. 17. The magnet losses are reduced by using step-skewed segmented magnets instead of a whole piece of continuous-skewed magnet. Fig. 18 shows the magnet eddy current loss density distributions comparisons between a continuous-skewed magnet and the step-skewed magnets. The 10-segmented step-skewed magnet reduces the magnet loss by over 90%. In order to evaluate the effect of skew on the magnet and iron losses of the unaligned stator-rotor configurations, 0-degree, 8- degree, 18-degree, 28-degree, and 38-degree stator-rotor skew angles are modeled and calculated by 3-D FEA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000872_iweca.2014.6845682-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000872_iweca.2014.6845682-Figure4-1.png", + "caption": "Figure 4. First mode shape of the assembly.", + "texts": [ + " The natural frequency comparison of order one is listed in table I. The FEA result of the single workpiece is extremely faithful to the EMA. It can be concluded that FEA is reasonable and precise. The natural frequency of the workpiece is found to be quite high. As can be observed in Fig. 3, notable mode shape is revealed in the part of cross thin wall. Therefore, it is the VSR equipment through the vibration platform that is capable of treating the workpiece. The mode shape of the assembly in the FEA is shown in Fig. 4. It should be pointed out that the VSR experimental application is based on the FEA results. The workpiece is placed on the strongly vibratory and easily clamped position, and three rubber supporters are placed on the nodal lines regions. In the FEA results, the natural frequency of the VSR equipment is 65.78 Hz, and then it increases to 68.05 Hz under the condition of assembly. However, those changes are hard to distinguish in the EMA results. Harmonic analysis is an effective means to carry out the steady state response to harmonic loads of a linear structure, and it is solved on the basis of modal analysis results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003855_ddcls49620.2020.9275206-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003855_ddcls49620.2020.9275206-Figure3-1.png", + "caption": "Fig. 3: The coordinate system of the UVMS", + "texts": [ + " 2\u03c0\u03c6 is the initial phase of the waves. When value of f is not 0, with the time t goes, a wave is considered propagating along the axis X . The force generated by this wave can be divided into two types: thrust T along the axis X and lateral force L along the axis Z. The force along the axis Y has little effect and is ignored. The training environment for the RL algorithm is based on the dynamic model of the UBVMS, here three DoF of kinematics and dynamics are considered. The coordinate system is shown in Fig. 3. To describe the position and posture change of the UBVMS, a fixed coordinate system E-xy and a moving coordinate systerm O-uv are established. The origin of system O-uv is fixed on the geometric centre of the robot. In the system E-xy, the pose of the robot is represented by a vector: \u03c7 = [x, y, \u03c8 ] T (2) where x and y denote displacements of the robot on the two axes. \u03c8 denotes angular displacement of the robot on horizontal plane. In the systemO-uv, the velocity of the robot is determined by a vector: v = [u, v, r ] T (3) where u and v denote linear velocity of the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001769_robio.2015.7418980-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001769_robio.2015.7418980-Figure7-1.png", + "caption": "Fig. 7 Calibration of TCP with two ball", + "texts": [ + " It requires controlling robot to make two pointed ends almost coincide. However, it is difficult to operate robot manually to achieve and the accuracy cannot be guaranteed. In order to solve above problems, a novel method of calibration is proposed in this paper. Instead of using pointed end, two balls are used in this method. One ball is installed in the place of drill and another one is fixed in experiment table. Therefore, the coordinates of the ball center which installed in end effector are coordinates of TCP. As is shown in Fig. 7, to get the coordinates of the ball center, operate the robot to make two ball touch from different ways. Owing to the radii of ball are equal in all directions, the distance between two balls is constant. Therefore, the eq.(17) and (18) are obtained. 2 2 2 2 1 2( ) ( ) ( ) ( )b m b m b mx x y y z z r r (17) 6[ 1] [ 1]T T b b b f f fx y z T x y z (18) Where xb, yb, zb are the coordinates of small ball in base coordinate. xm, ym, zm are the coordinates of big ball in base coordinate. r1, r2 are radius of two balls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003580_j.surfcoat.2020.126478-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003580_j.surfcoat.2020.126478-Figure1-1.png", + "caption": "Fig. 1. 3D and contour plots of the effect of pulse width, scanning speed and operational distance of the pulse laser beam on microhardness of composite areas (a and b) pulse width of 6 (ms), (c and d) pulse width of 8 (ms), (e and f) pulse width of 10 (ms).", + "texts": [ + " = + + + + = = = = Y x x x x i k i i i k ii i i k j k ij i j0 1 1 2 1 2 (4) The mathematical equation derived for best- fitting of the microhardness value in terms of input data in the processed region is given as stated in Eq. (5). Given the results in Table 5, the accuracy of the proposed relationship can be clearly assessed and confirmed. X X 27 X 12 5 X 10 X 1 2 3 1 2 1 3 2 3 1 2 2 2 3 2 = + + + + + ( ) . . . (5) To investigate the accuracy of the proposed model, it is vital to evaluate R2 and p-values. In this regard, according to Table 5, the R2 value of 0.987 demonstrates good accordance between experimental and predicted data with and a p-value of 0.001 confirming statistical significance. Fig. 1 illustrates the effects of variation in scanning speed, pulse width, and laser beam working distance on the microhardness of the processed region. According to Fig. 1, it can be stated that, by decreasing laser beam scanning speed, more thermal energy is introduced into the work piece which increase the temperature of the molten pond, thereby increasing the dissolved TiC content tends to happens. On the other hand during cooling, new complex carbides are formed consisting of a combination of TiC and the matrix phase alloying elements, therefore lower hardness is achieved when compared to the original (dissolved) TiC phase. As a result, the processed area hardness decreases with decreasing laser beam scanning speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003969_0954406220983367-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003969_0954406220983367-Figure7-1.png", + "caption": "Figure 7. Picture of linear guideway (a) linear guideway of machine tool; (b) contact surface of linear guideway.", + "texts": [ + " As for the real contact area calculation, the proposed method based on fractal theory is compared with the Finite element method, former fractal method and GWmethod. Results dem- onstrate that the proposed method improves the modeling accuracy of real contact area of linear guideway. In this section, the flow chart of verification of proposed method is shown in Figure 6. Friction and wear usually occur in the contact surfaces of slider, guideway and roller for a long-term motion of machine tool, therefore it is crucial to measure and analyze the interface condition of linear guideway in Figure 7. The microstructure of the contact surface topography is measured by laser micro- scope 3d & profile measurement instrument from Keyence type VK X150, whose display resolution and spot diameter of laser light are in turn 5nm and 0.4lm. Generally, slider and guideway interfaces are cut to small experiment samples. The sample length of slider and guideway are both 20 10 3mm, and the con- tact surface are all cleaned with acetone and dried. The topography measurements of the contact surface of slider and guideway are exhibited in Figure 8(a) to (c), respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003711_j.eml.2020.101076-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003711_j.eml.2020.101076-Figure3-1.png", + "caption": "Fig. 3. Completely rigid body model of CBR.", + "texts": [ + " the left leg and the right leg are both in contacting with the slope. When the leg is swinging, it means the leg has no contact with the slope. And it will rotate with respect to the pin of hip. When the leg contacts the slope, the transient stress waves will be generated by the oblique impact and propagate from the foot of the support leg to the rest components of the robots. 2.1. Completely rigid body model (CRBM) Neglect structural deformation and wave propagation effect, the CRBM discussed in this paper is shown in Fig. 3. For analyzing more conveniently, in this section the CRBM is simplified into a planar model. The leg is assumed to be a rigid rod and modeled as a line segment (i.e. the cross-sectional area is zero) with inertia. The COF between the leg and the pin is ignored. Meanwhile, the slope is modeled as a rigid ground. The connection at the hip is modeled as a plane joint and the mass of the pin is modeled as a concentrated mass attached on the pin. The mass of the pin and the leg are mp = \u03c1p \u00d7\u03c0 \u00d7 r2 \u00d7h and m = \u03c1l \u00d7 l\u00d7A, respectively", + " l p a o s s i 2 ( I g w o p o x n a i l s c t w s t a M w p H ( w f Since the phases 1 and 3 in Fig. 2 have the same contact state nd the phases 2 and 4 have the same contact state, the walking f the CRBM is further classified into two phases, which are (1) ingle-leg support phase, i.e. support leg is in contacting with the lope, swing leg is swinging; and (2) double-leg support phase, .e. two legs are both in contacting with the slope. .1.1. Dynamic equations of CRBM 1) Dynamic equations during single-leg support phase Fig. 3 shows the CRBM is during the single-leg support phase. t is assumed that there is no slip at the contact point. The eneralized coordinates of the CRBM system are q = [q1 q2] T, here q1 = [x y]T and q2 = [\u03b81 \u03b82] T. (x, y) is the coordinates f the contact point in inertia system x-o-y. Down the slope is the ositive direction of the coordinate axis x, and the exterior normal f the slope is the positive direction of the coordinate axis y. Let = y = 0 when time t = 0. \u03b81 and \u03b82 are the angles from the ormal line (i.e. the short dash line in Fig. 3) of the slope to the xis of the support leg and the swing leg, respectively. The angle s positive for counterclockwise as rotating the normal line to the eg axis. Hence, it has \u03b81 < 0 and \u03b82 > 0 for the configuration as hown in Fig. 3. It should be noted that when |\u03b81| > |\u03b82|, the y oordinate of the swing foot is less than zero, which means that he foot of the swing leg is lower than the slope. Hence, the foot ill penetrate into the slope. During the single-leg support phase, the system only has conervative force (i.e. gravity) to do work. Based on the conservaion law of mechanical energy, the second Lagrange equation is pplied to derive the dynamic equations of the CRBM as (q)q\u0308 + C (q, q\u0307) + G(q) = W (q)Tf (1) here M(q) = \u22022T \u2202 q\u03072 , G(q) = ( \u2202V \u2202q )T , f = [ft fn]T C (q,q\u0307) = [ \u2202 \u2202q (( \u2202T \u2202 q\u0307 )T )] q\u0307 \u2212 ( \u2202T \u2202q )T , W (q) = dr dq T indicates the kinetic energy of the system, V indicates the potential energy of the system, M(q) is called the matrix of inertia, vector G(q) is gravitational forces, C (q, q\u0307) is the vector of velocity-dependent terms such as Coriolis and centripetal forces, W (q) is the robot\u2019s Jacobian with respect to the support foot\u2019s osition and f is the normal and tangential components of the contact force acting at the support foot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003656_s10015-020-00655-x-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003656_s10015-020-00655-x-Figure1-1.png", + "caption": "Fig. 1 Overview of the proposed UAV robot", + "texts": [ + " After constructing a simplified dynamical model of the UAV robot, a computed torque controller is designed for the UAV robot and a coordinate transformation is used to simply the allocation problem. Some simulation results of the proposed UAV robot are given to demonstrate the tracking performance of arbitrary position and attitude. A simulation with wind disturbances is also added to verify the position control performance of the proposed control strategy in the final part. The overview of a prototype UAV robot developed in Xu et\u00a0al. [9] is shown in Fig.\u00a01. The definition of the coordinate systems of the prototype UAV robot is shown in Fig.\u00a02. The proposed UAV robot can be considered as a multibody system including three rigid bodies, i.e., two tiltable coaxial rotors Pi and a main body B . In each tiltable coaxial rotor, two RC servomotors are set for constructing a 360\u25e6 tilt mechanism, and one RC servomotor is also set for making another 180\u25e6 tilt mechanism, while two brushless motors are prepared for actuating a pair of DJI 1038 propellers. In the main body, it includes an aluminum pipe structured for the body frame, four passive wheels that are used on a wall or floor, four guide wheels that are available in a transition mode from a floor to a wall (or in its reverse transition mode), a battery and all other electronic devices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000034_roma46407.2018.8986730-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000034_roma46407.2018.8986730-Figure4-1.png", + "caption": "Figure 4: Model of the slide potentiometer The slide potentiometer is calibrated according to the position of the fingers. The values of the accelerometer output at every position is given in the below table.", + "texts": [], + "surrounding_texts": [ + "The system consists of two sides one is the sending side and the other is the receiving side. The data is transmitted wirelessly through the Zig-bee protocol. The overall block diagram of the system is shown in figure 1." + ] + }, + { + "image_filename": "designv11_34_0000388_romoco.2015.7219734-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000388_romoco.2015.7219734-Figure13-1.png", + "caption": "Figure 13: Schematic of notation", + "texts": [ + ", in body frame the control torque is not restricted to a linear subspace. With this in mind, the configuration space of the MAV viewed as a mechanical system is Q = R3\u00d7SO(3)\u00d7Tp with R3 \u00d7 SO(3) corresponding to the configuration space of the main body (3D position and orientation), and Tp corresponding to the configuration space of propellers and moving surfaces in the main body\u2019s frame (i.e., parametrization of propellers\u2019 rotation axes, rotation angles of propellers and moving surfaces). The notation for the configuration variables is now defined (see Fig. 13). \u2022 I = {O; i0, j0,k0} denotes a fixed inertial frame with respect to (w.r.t.) which the vehicle\u2019s absolute pose is measured. This frame is chosen as the NED frame (North-East-Down) with i0 pointing to the North, j0 pointing to the East, and k0 pointing to the center of the Earth. \u2022 B = {G; i, j,k} denotes a frame attached to the main body, with G the vehicle\u2019s center of mass. It is assumed here that G is a fixed point in the body frame. This may not always be true, especially for tilt-rotors/tilt-wings MAVs since their shape can change, but this is a reasonable assumption in first approximation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003155_1350650120928661-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003155_1350650120928661-Figure4-1.png", + "caption": "Figure 4. Schematic diagram of surface stress calculation.", + "texts": [ + ",43,44 surface stress and shear stress of the workpiece surface can be obtained under a given load. Similar to the work of others,21\u201328 for the convenience and efficiency of the calculation of contact model, the interaction between asperities is not taken into account. In virtue of the elastic half space theory and the contact mechanics knowledge put forward by Johnson,45 it is easy to calculate the value of surface stress field including rmax and smax, where the unit of roughness parameter is micrometer and the unit of stress is megapascal (Figure 4). Here are the surface stress field calculation formulas x \u00bc P 2 1 2 \u00f0 \u00de r2 1 z x2 y2 r2 \u00fe zy2 3 3zx2 5 \u00f01\u00de y \u00bc P 2 1 2 \u00f0 \u00de r2 1 z y2 x2 r2 \u00fe zx2 3 3zy2 5 \u00f02\u00de z \u00bc 3P 2 z3 5 \u00f03\u00de xy \u00bc P 2 1 2 \u00f0 \u00de r2 1 z xy r2 xyz 3 3xyz 5 \u00f04\u00de xz \u00bc 3P 2 xz2 5 \u00f05\u00de yz \u00bc 3P 2 yz2 5 \u00f06\u00de where \u00bc x2 \u00fe y2 \u00fe z2 1=2 , r2 \u00bc x2 \u00fe y2. The total area of the sampling area is: An\u00bcLx Ly\u00bc 0.4 0.56\u00bc 0.224mm2. The calculation method in this paper is compared with the numerical method results in Wang et al.46 Under 1N load, based on the surface topography of the workpiece shown in Figure 5, the surface stress calculation results of the two methods are shown in Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003430_lra.2020.3015463-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003430_lra.2020.3015463-Figure6-1.png", + "caption": "Fig. 6. Rotational moment obtained from the wall reaction force in three cases: (a) Front sub-tracks are kept flat, (b) front sub-tracks are lowered, and (c) front sub-tracks are raised. the distance between the collision point and the robot\u2019s center is maximized in (a), so that the largest rotational moment is obtained if the wall reaction force is the same.", + "texts": [ + " Based on the results, we set the same velocity for the left and right main-tracks to use the wall reaction force to climb up the spiral stairs. Motion of sub-tracks in \u201cClimbing\u201d phase in Fig. 3 is discussed considering how to keep the robot\u2019s posture and how to obtain rotational moment using the wall reaction force. Here, the flipper control in \u201cClimbing\u201d phase is discussed from the perspective of 1) obtaining rotational moment and 2) preventing catches with the wall or the ground. 1) Obtaining Rotational Moment: Fig. 6 shows rotational moment obtained from the wall reaction force. This figure considers three cases: (a) sub-tracks are kept flat, (b) sub-tracks are lowered, and (c) sub-tracks are raised. The distance between the collision point and the robot\u2019s center is maximized in (a), so that the largest rotational moment is obtained if the wall reaction force is the same. Thus, (a) keeping sub-tracks flat is the best choice in terms of rotational moment by wall reaction force. 2) Preventing Catches With the Wall or the Ground: Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002419_ijmr.2016.079461-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002419_ijmr.2016.079461-Figure2-1.png", + "caption": "Figure 2 Phographic view of the customised bearing test rig (see online version for colours)", + "texts": [ + ", normal (N), bearing with defect on ball (B), bearing with defect on inner race (IR) and bearing with defect on outer race (OR) at varying load and speeds has been considered for assessment. Raw vibration signals have been denoised using the Interval-dependent scheme, selected from among five schemes based on kurtosis and RMSE. The denoised signals are decomposed using discrete wavelet transform (DWT) and 17 statistical features have been extracted from the second level detailed wavelet coefficients (cD2). ANN has been used to evaluate PCA-based DRTs. The proposed bearing fault diagnostic methodology is depicted as shown in Figure 1. Figure 2 shows the photographic view of the customised bearing test rig used for acquiring vibration signals. It has a shaft mounted on support bearing and test bearing. The shaft is driven by an induction motor with variable speed regulator. The drive motor and bearing test rig were driven through a timer belt pulley arrangement, with a speed ratio of 2.25. Vibration signals were acquired at variable load (0, 1.7 kN) and variable speed conditions (222 rpm and 622 rpm). In the present work, 6,205 deep groove ball bearing is considered for testing. A conical shape defect (3 mm diameter \u00d7 2 mm depth) is machined on the inner race, ball and outer race elements of the bearing. Radial load on the test bearing is applied using a hand pump set-up. Two accelerometers are placed on the housing of test bearing as shown in Figure 2. The accelerometers are coupled to signal conditioning amplifiers and in turn to data acquisition system (DAQ) hard ware installed in the computer through connecting cables. A customised LABVIEW (vi) programme is used for acquiring the vibration signals at 48 kilo samples/s and is saved in the computer as .txt file. In this work, acceleration signals at the specified sampling rate are collected for 5.08 seconds. Signals acquired from vertical accelerometer (accelerometer-X) are used for analysis, as the signal acquired in horizontal accelerometer (accelerometer-Y) is found to be not very sensitive to bearing conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001027_detc2013-13712-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001027_detc2013-13712-Figure4-1.png", + "caption": "Figure 4. Four pad tilting-pad bearing used in analysis", + "texts": [ + " (12), the new positions and velocities are calculated, which are used to update the bearing forces at each time step. Figure 3 is a flowchart illustrating the calculation of the transient analysis. A 4-pad tilting-pad bearing is numerically tested for transient analysis first and used in following rotor-bearing systems for time transient analysis. The tilting-pad bearing information used in the model is shown in Table 1, and the geometry of the bearing including shaft, pad/pivot position and bearing clearance, is shown in Fig. 4. In the figure, the clearances have been exaggerated for illustration purposes. Table 1 Tilting-pad Bearing information Type: Tiling Pad Bearing, Load Between Pads Property Value Unit Pad Angle 5 \u03c0 /12 rad Pivot Location \u03c0/4, 3 \u03c0 /4, 5 \u03c0 /4, 7 \u03c0 /4 rad Offset 50% Length 0.1524 m Diameter 0.1524 m Preload 0 Bearing Clearance 1.524\u00b710-4 m Pad Clearance 1.524\u00b710-4 m Viscosity 1.15\u00b710-2 Pa.s Pivot stiffness 8.756\u00b7108 N/m Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/16/2016 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure3.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure3.6-1.png", + "caption": "Fig. 3.6 Magnetic field within a brushless exciter during steady state operation coupled to the turbo-generator field winding", + "texts": [ + " The remaining 9 loops lead via mere diode paths and the load: ( )( ) 0371k2l121k2l12 =++ ++\u22c5+\u22c5+\u22c5+\u22c5 ht uuu (3.37) for k = 1,2,3 and l = 0,1,2. In comparison with this integrated network solution of the FD calculation an iterative solution is described in [3.25], where the currents in each winding are estimated, before the field in the machine is solved. This approach is only feasible for machines with concentrated windings. A verification of the mathematical model with measurements is realized for a brushless exciter acc. fig. 3.6. Extensive investigations have been done on this 8-pole brushless exciter with rotating armature. The diode currents have been measured by Rogowski coils. The signals were transmitted by telemetry. 3.2 Non-linear Transient Time-Stepping Numerical Field Calculation 43 The exciter has been measured on an exciter test bench. Therefore the rectified current has been fed in a 0.1 \u03a9 resistor standing for the main turbo-generator field winding. This has been considered in the calculation in the same way" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000856_tmag.2014.2362515-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000856_tmag.2014.2362515-Figure3-1.png", + "caption": "Fig. 3. Analysis model.", + "texts": [ + " In this paper, the amplitude and angle of the principal stresses are approximated as \u03c31 = \u03c3x , \u03c32 = \u03c3y, \u03b8\u03c3 = 0\u00b0. We analyzed the magnetic characteristic distributions in the motor model with the SCES modeling. The governing equation of the 2-D magnetic characteristic analysis considering of the SCES modeling can be expressed as \u2202 \u2202x ( \u03bd\u0304yr \u2202 A\u0307 \u2202x + j\u03c9\u03bd\u0304yi \u2202 A\u0307 \u2202x ) + \u2202 \u2202y ( \u03bd\u0304xr \u2202 A\u0307 \u2202y + j\u03c9\u03bd\u0304xi \u2202 A\u0307 \u2202y ) = \u2212J\u03070z (3) where Az is the magnetic vector potential, and J0 is the exciting current density. Fig. 3 shows the permanent magnet motor model used in this numerical simulation. The residual magnetization is 0.2 T, and exciting current is 5 A, and the speed of revolution is 1500 r/min. A non-oriented electrical steel sheet is used as the magnetic material of the stator and rotor cores. We try to reduce the magnetic power loss of the motor core by applying the tensile stress. The uniaxial tensile stress is applied to the region, which is magnetized under the alternating magnetic flux condition. The biaxial tensile stress is applied to the region, which is magnetized under the rotating magnetic flux condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001554_10402004.2013.856979-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001554_10402004.2013.856979-Figure1-1.png", + "caption": "Fig. 1\u2014Rolling four-ball test setup (color figure available online).", + "texts": [ + "25 stress cycles on the driving ball running track. This ratio has been found to hold over a wide range of conditions, even in the EHL regime (Ku, et al. (23)). Test parameters used in this study are similar to those developed by Hamaguchi, et al. (22), which reportedly provided good repeatability and reasonable correlation to other gear fatigue test methods in regard to lubricant-dependent pitting results. Tests were performed using a computer-controlled Phoenix TE 92 HS four-ball test machine in rolling test configuration, as shown in Fig. 1. Once the test commences, the raceway holder containing the three planet balls is loaded against the driving ball with the specified load. This is achieved by means of a pneumatic load actuator. A heating pad below the lower assembly ensures that the specified bulk oil temperature is maintained. Prior to each test, the holder including the raceway, four new balls, as well as the collet were cleaned in heptane for 8 min in an ultrasonic cleaner. Thereafter the parts were rinsed in ethanol and dried" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002451_cacs.2015.7378369-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002451_cacs.2015.7378369-Figure6-1.png", + "caption": "Figure 6. Illustration of current reconstruction problem at a low-modulation index: (a) Reference voltage vector in Sector I boundary region. (b) Corresponding three-phase PWM pattern. ( c) Voltage vectors for switching (proposed). (d) Corresponding three-phase PWM pattern (proposed).", + "texts": [ + " The same technique also ensures the PWM output signals are divided into two parts equally, and keep the waveform symmetrical in Fig. Sed). Furthermore, the add-on voltage vector V4' ensures the current measurement is valid during the two switching states and the sampling can be completed at the left side within one PWM interval. C. In the Case of Low Modulation Index (P3) At a low modulation index, the duty cycles of the PWM signals for the three phases will have an almost equal duration. The result is that the active voltage vector is not being used for long enough to ensure a proper sampling of the DC-link current. Fig. 6 illustrates the problem in a double-sided modulation strategy. Thus, two currents cannot be sampled correctly because V2 and V; are shorter than Tmm . It is necessary to generate a minimum time delay Tmin from which a proper sampling can be obtained. Taking both \ufffd and T2 as less than 2Trrun in sector I for example, both \ufffd and T2 are replaced by 2Tmin to ensure the phase current reconstruction detection demand is satisfied. Then, two complementary voltage vectors, V4' and Vs' are applied during \ufffd = 1'1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000479_3dp.2015.0011-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000479_3dp.2015.0011-Figure2-1.png", + "caption": "Figure 2. Tensioning system and force arrow with Jaws filament drive bolt.", + "texts": [ + " We suspect the reliability of the filament feed gear is dependent upon three factors: (1) size of contact surface between the drive gear and the filament; (2) depth of the gear \u2019 s tooth engagement into the filament; (3) number of teeth engaged in the filament at any one time; and (4) the direction of the force vector imparted by the filament drive gear into the filament. R&D at re:3D has delved into this problem and below is the result of their work. The goal of this work is to characterize the amount of force that can be generated\u00a0 by a machined extruder drive\u00a0 bolt, affectionately named Jaws.\u00a0 Figure 1 shows a 3D rendering of\u00a0 the filament drive gear, Jaws, that will\u00a0be mounted to the shaft of a geared NEMA 17 stepper motor. The testing outlined below was performed with a Greg \u2019 s Wade \u2013 type extruder for 3\u00a0 mm polylactic acid (PLA) filament (see Fig.\u00a02). The Jaws filament drive gear is machined using a four - axis computerized numeric control milling machine. In the design process, our aim was to optimize the four variables stated above. \u2022 The amount of contact between the drive gear and the filament is maximized by machining the gear teeth using a cutting tool of the same diameter as the filament it will drive. In our case, the filament is 3.0\u00a0mm in diameter. \u2022 The depth of tooth engagement was optimized by balancing greater tooth engagement against the number of teeth engaged into the filament at any one time", + " By studying the section view in Figure 3 you can see the depth of engagement, the number of teeth engaged, and the force vector for a 25 - tooth drive gear. You can view a short video of the machining process on our re:Tech YouTube channel. The first objective of this study is to determine the optimal tension setting\u00a0 for\u00a0 the Greg \u2019 s Wade extruder. Tension is\u00a0 adjusted by rotating a pair of\u00a0 screws that compresses a pair of springs, which in turn presses the extruder \u2019 s idler bearing against the filament (see Fig.\u00a02). Figure 3. Tooth engagement, number of teeth engaged, and force vector. 3D PRINTING 87 VOL. 2 NO. 2 \u2022 2015 \u2022 DOI: 10.1089/3dp.2015.0011 If the tensioner is adjusted too loosely, the filament will not fully engage against the teeth of the drive bolt. If the tensioner is adjusted too tightly, there may be excessive friction and wear in the extruder. To accurately measure the force imparted to the filament by the extruder we felt it was important to measure the ability of the extruder to push the filament down into the hot - end", + " The load cell was connected to a load cell amplifier (TMO - 1 from Transducer Techniques), and the analog output from the amplifier was measured by a 12 - bit A to D daq (USB - 1208LS from Measurement Computing). Data was collected at 100 Hz and saved to csv file format. The daq was calibrated using a\u00a0 mass of solid brass and aluminum of\u00a0 known volume. Collected data was processed and graphed using custom Matlab code. The extruder is driven with a 1.68 A 72 oz - in NEMA 17 stepper motor powered at 24 volts. The test setup can be seen in Figure 2 above. All testing was done with a hot - end temperature of 210 \u00b0 C. Three trials were performed at four different levels of tensioner adjustments. Tensioner adjustment levels were determined as 2, 4, 8, and 12 revolutions (1 revolution = 360 degrees of rotation) of the tensioner adjustment screw beyond full engagement into its corresponding nut. Example trials are shown below. Notice in the data graph the period of zero force at the beginning of the trial is followed by a gradual increase of force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002346_chicc.2016.7554341-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002346_chicc.2016.7554341-Figure6-1.png", + "caption": "Fig. 6 Kinematic model of Fig. 8 Leader-follower model car-like mobile robot and state error", + "texts": [ + " 2 depicts the state update strategy of differential-drive mobile robot and the process can be described as equation (25) ~ (29). (25) i d i i d i V i x u K C p y Then we have robot target speed as cos sin sin cos 1 (26) cos cos sin cos d d l d d r tt t d d xb b d dr y b b Considering the limitation of robot\u2019s speed, the robot real speed can be decided as ( ) ( ) ( ) ( )max ( ) ( )max ( ) ( )max (27) d d l r l r l r l r d l r l r l r l r if if And the sensor\u2019s real speed can be written as 4.2 Car-like mobile robot Fig. 6 shows the kinematic model of a car-like mobile robot and its kinematic equation can be denoted as equation (30). The follows are the notations used C : basic position of car-like mobile robot L : length between front axle and rear axle : angle between front wheel and vehicle body v : velocity of mobile robot : speed of front wheel rotation Specially, in order to simplify the calculation, it is assumed that the measurement of C is accurate. Fig. 7 describes the solution to the state update law which drives a car-like robot to its target state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000761_ivs.2013.6629515-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000761_ivs.2013.6629515-Figure9-1.png", + "caption": "Fig. 9. Trajectories of the robots attaining the formation.", + "texts": [ + " Experimentations are made on Khepera III robots and illustrate a navigation in formation of the robots while avoiding each other. A central camera, at the top of the platform gives positions of all the robots and the obstacles thanks to circular bar codes installed on them. The objective, in short term horizon, is to use the local sensors of the robots in order to get a completely decentralized architecture. The scenario illustrates three robots which have to join a triangular virtual structure. The latter moves along a circular trajectory (cf. Figure 9). Robots are put in their initial conditions so that they must avoid each other before joining the formation. It is observed that the collision avoidance is successfully accomplished for all the robots. Moreover, no conflict was observed since avoidance is done in one direction (robots of the same system)(cf. Section II-C.3). The formation is attained as shown in figure 9 illustrating the trajectories of the three robots. The proposed control architecture devoted to the navigation in formation in presence of obstacles must be enriched to generate only attainable set-points. In fact, the proposed control law is theoretically stable. However, in practice, additional constraints must be taken into account. In our case, kinematic constraints (maximum velocities) of the robot imposes to define maximum authorized set-points. It is then proposed to study the obstacle avoidance controller case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003349_tmag.2020.3011612-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003349_tmag.2020.3011612-Figure5-1.png", + "caption": "Fig. 5. Cross section of the 12 stator slots, 10 poles FSCW and 16 rotor bars.", + "texts": [ + " 3 0 0 0 0 0 0 bar bar bar end end end n R R R R R R R (13) , , , , 1 , 1, 3 0 0 0 0 0 0 0 0 0 0 i j i n i i j n n i n n n n n L M M M M L M M L (14) To validate the proposed approach (the equivalent electrical circuit and the calculation of its components), two different squirrel cage induction machines are studied. The first one is a distributed winding (DW) with 48 stator slots and 30 rotor bars as shown in Fig. 4. The second one is a fractional slot concentrated winding (FSCW) with 12 stator slots and 16 rotor bars as shown in Fig. 5. As the second winding is a double layer one, each stator slot houses two coil sets of either the same phase or two different phases. For each machines results are given for two operating points i.e. at synchronous speed (no load) and at load (slip s = 5%) and compared to the ones obtained by a 2D FEM in the same conditions while considering a linear B(H) curve of the magnetic material. A. Results for the DW machine Fig. 6.a shows the variation of the self-inductance of the first bar, L1, and Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003756_1350650120966894-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003756_1350650120966894-Figure1-1.png", + "caption": "Figure 1. Contact path of two tooth surface under ideal condition.", + "texts": [ + " He pointed out when a pair of teeth are meshed near the pitch point, the sliding friction coefficient should be close to 0 because their relative sliding velocity tends to 0, which is close to pure rolling. However, results from the current calculation formulas of the sliding friction coefficient are generally large. Helical gear is widely used in high speed and heavy load transmission because of its high load capacity. In ideal condition, the contact between two teeth surface is line and the change of length of contact line is short-long-short (Figure 1). However, considering machining error/installation error/modification of gear, the contact between two teeth surface is point, not line,14 which is shown in Figure 2. Under load, the contact point becomes contact ellipse. The major axis of ellipse is much larger than the minor axis of ellipse and therefore it is still taken as contact line. Figure 3 shows the projection of contact line on the middle section under load.15 But the line contact considering machining error/installation error/ modification of gear is completely different with that in ideal condition and can\u2019t be dealt with in the traditional method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001194_s11370-015-0170-5-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001194_s11370-015-0170-5-Figure1-1.png", + "caption": "Fig. 1 A differential-drive robot", + "texts": [ + " In contrast to B-spline curves, the Hermite curves can take into account the orientation of an AMR at each via-point, thereby making it possible to apply DP to our problem. For all paths generated in the proposed DP algorithm, the trajectory planner generates the time-optimal trajectory considering the kinematic and dynamic constraints. Section 2 defines the motion planning problem, and Sect. 3 describes the proposed motion planning method. Section 4 describes our simulation results for an AMR model and verifies the usefulness of the proposed method. We give our conclusions in Sect. 5. Figure 1 shows a differential-drive robot with two wheels which is widely used for AMRs. Let (x, y, \u03b8) be the position and orientation of the robot in Cartesian space, and (\u03d51, \u03d52) be the position of the right and left wheel in joint space. The forward kinematic equation of the robot is given as [1]: x\u0307 = r 2 (\u03d5\u03071 + \u03d5\u03072) cos \u03b8 (1) y\u0307 = r 2 (\u03d5\u03071 + \u03d5\u03072) sin \u03b8 (2) \u03b8\u0307 = r 2L (\u03d5\u03071 \u2212 \u03d5\u03072) (3) where r is the diameter of the wheel and L is the distance from the robot\u2019s center to the wheel. The speed of each wheel should be constrained by its limit such that |\u03d5\u03071| \u2264 \u03d5\u03071,max (4) |\u03d5\u03072| \u2264 \u03d5\u03072,max (5) Let (\u03c41, \u03c42) be the torque of right and left wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003855_ddcls49620.2020.9275206-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003855_ddcls49620.2020.9275206-Figure1-1.png", + "caption": "Fig. 1: The design of the UBVMS", + "texts": [ + " In Section 5, a reinforcement learning method based on DDPG theorem is presented to solve the MDP. In Section 6, the training result of the reinforcement learning algorithm is presented. The position control is implemented in 5 cases and the simulation results are analyzed. The position control is implemented on an UBVMS. Some relevant equipments and materials are introduced in this section. Authorized licensed use limited to: East Carolina University. Downloaded on June 15,2021 at 19:20:41 UTC from IEEE Xplore. Restrictions apply. The prototype design of the UBVMS is shown in fig. 1. It possess two biomimetic propellers to achieve multimodal motions, a 5-DoF (degree of freedom) manipulator for underwater operation and two pairs of binocular cameras for underwater detection and observation. Some actual measured parameters of the UBVMS are listed in the Table 1. The biominetic underwater propeller possesses a flexible undulating fin which consists of twelve short rays connected by a black silicone sheet. By distributing all rays in interval phases of a sinusoid, propulsive force can be generated by the propeller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.33-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.33-1.png", + "caption": "Figure 9.33 A Puma manipulator having the fourth joint fixed.", + "texts": [ + "548) \ud835\udf14\u2217 3 = 0 (9.549) Kinematic Analyses of Typical Serial Manipulators 303 Equation (9.549) constitutes the task-space restriction of this version. In this version, the joint velocities s\u03072, s\u03074, and ?\u0307?5 are still expressed in the same way as in the first version, but ?\u0307?3 and ?\u0307?6 assume the following different expressions. ?\u0307?3 = \ud835\udf0e\u20323w3\u2215s4 (9.550) ?\u0307?6 = \ud835\udf0e\u20325(\ud835\udf14 \u2217 2 \u2212 ?\u0307?3) = \ud835\udf0e\u20325(s4\ud835\udf14 \u2217 2 \u2212 \ud835\udf0e \u2032 3w3)\u2215s4 (9.551) As an example of a deficient manipulator, consider the special Puma manipulator shown in Figure 9.33, which is used deficiently, either deliberately or due to an actuator defect, so that the fourth joint remains fixed with \ud835\udf034 = 0. Therefore, instead of studying as a new manipulator, this deficient manipulator is studied here with the necessary relevant modifications made in Section 9.1, where a regular Puma manipulator is studied. (a) Joint Variables \ud835\udf031, \ud835\udf032, \ud835\udf033, \ud835\udf035, \ud835\udf036 Note that \ud835\udf034 = 0 because the fourth joint is fixed. (b) Twist Angles \ud835\udefd1 = 0, \ud835\udefd2 = \u2212\ud835\udf0b\u22152, \ud835\udefd3 = 0, \ud835\udefd4 = \ud835\udf0b\u22152, \ud835\udefd5 = \u2212\ud835\udf0b\u22152, \ud835\udefd6 = \ud835\udf0b\u22152 (c) Offsets d1 = 0, d2 = OS, d3 = 0, d4 = ER, d5 = 0, d6 = RP (d) Effective Link Lengths b0 = 0, b1 = 0, b2 = SE, b3 = 0, b4 = 0, b5 = 0 (e) Link Frame Origins O0 = O,O1 = O,O2 = S,O3 = E,O4 = R,O5 = R,O6 = P 304 Kinematics of General Spatial Mechanical Systems (a) Link-to-Link Orientation Matrices C\u0302(0,1) = eu\u03031\ud835\udefd1 eu\u03033\ud835\udf031 = eu\u03033\ud835\udf031 (9", + "605), it constitutes a constraint on the tip point velocity so that only one of the components v1 and v2, say v1, can be specified freely as desired. The other one must satisfy the following velocity constraint equation in the task space. v2 = v1 tan \ud835\udf031 + (d6s\ud835\udf03235 + d4s\ud835\udf0323 + b2c\ud835\udf032)?\u0307?1 sec \ud835\udf031 (9.619) In this case, the manipulator may have two kinds of motion singularities, which are described and discussed below. (a) First Kind of Motion Singularity Equation (9.607) implies that the first kind of motion singularity occurs if p\u2032 1 = d6s\ud835\udf03235 + d4s\ud835\udf0323 + b2c\ud835\udf032 = 0 (9.620) On the other hand, referring to Figure 9.33 and Part (d) of Section 9.7.2, it is seen that p\u2032 1 is the projection of the vector r\u20d7SP on the first axis of the link frame 1(O). Therefore, in this singularity, r\u20d7SP becomes parallel to the axis of the first joint. In other words, the tip point P becomes located on a cylindrical surface, whose base radius is d2. This singularity is illustrated in Figure 9.38. If the manipulator has to pass through this singularity or keep on moving on the singularity surface as a task requirement, then, according to Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000855_kem.579-580.300-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000855_kem.579-580.300-Figure1-1.png", + "caption": "Fig. 1 The generating of cylindrical gear", + "texts": [ + " (2), the condition of pitch curve without concave is obtained as follows. ))1/(1),1/(1( 2 2 22 1 2 \u2212\u2212\u2264 mnmnmink . (3) Modeling of Cylindrical Gear. The tooth profile of non-circular gear are enveloped by the tooth profile of cylindrical gear, so the tooth profile of cylindrical gear should has higher accuracy. This paper fit the points on the tooth profile with cubic spline curve, and scan a straight or helical tooth based on the generated involute profile, then array all teeth, the generated involute profile and cylindrical gear are shown in Fig. 1. Realization of Pure Rolling. As shown in Fig. 2, make four-order elliptic gear with concave as an example, coordinate system )-( xyoS is fixed with gear billet and )-( 111 yxoS is mobile with cylindrical gear but without rotation. When the pith circle roll from A to B , the turned \u03d5 of cylindrical gear around its center consists of the appendant angle 1 \u03d5 caused by the rotation around the center of polar coordinates and the angle 2 \u03d5 caused by the pure rolling arc length AB S . The coordinates of B point in coordinate system )-( 111 yxoS is as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001636_1.4032400-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001636_1.4032400-Figure3-1.png", + "caption": "Fig. 3 The coordinate system of the gear", + "texts": [ + " The method consists of four parts, including the gear tooth surface equations, the worm tooth surface equations, the coordinate transformation matrix between coordinate systems, and the algorithm for minimizing the distance between two entities. 3.1 Gear Tooth Surface Equations. There are many methods to establish the equations of the gear tooth surface [23]. In order to simplify the followed calculation, a special coordinate system is selected. The gear tooth surface is represented in the coordinate system Sg Og : Xg;Yg;Zg , as shown in Fig. 3. The Yg- axis is along the axis of the gear. The Zg-axis goes through the pitch point of the tooth surface. And the Xg-axis is determined by the right-hand law of Cartesian coordinate system. The origin point is determined by the intersection of the Yg-axis and the middle transverse plane of the gear. Assuming the rotational angle u of gear is zero when the real contact point is on the Zg-axis. The purpose of this assumption is to define the origin of abscissa of the GIE unit curve; thus, the positions of the start and end points of the involute stage on the theoretical GIE unit curve are known", + " The dashed curve AB denotes another transverse involute profile through an arbitrary point A, and point B is the start point of this involute profile. The Yg-axis coordinate of the plane that contains the profile AB is y. The pressure angle at point A of the transverse profile AB is a. So, the point A is determined by y; a\u00f0 \u00de. The value range of y is b=2; b=2\u00bd and the value range of a is 0; ata2\u00bd , where b is the face width and ata2 is the pressure angle of the tip of gear in transverse section. To simulate the process of meshing, it is necessary to establish the equations of the tooth surface after the gear rotating through an angle u. Figure 3 shows the initial position of the tooth surface where u \u00bc 0. After rotating an angle u measured from the initial position, the point A is determined by y; a;u\u00f0 \u00de. The value of u varies during the meshing process. The involute function is defined as inv a\u00f0 \u00de \u00bc tan a\u00f0 \u00de a (2) The helix angle of the gear is denoted by b2 ; then, the helix angle of the base cylinder of gear is bb2 \u00bc arctan r2=rb2 tan b2\u00f0 \u00de (3) The angle between the plane XgOgYg and the plane AOgYg is described as /AOX y; a;u\u00f0 \u00de \u00bc tan a\u00f0 \u00de inv atw2\u00f0 \u00de \u00fe y tan bb2\u00f0 \u00de=rb2 \u00fe u (4) 033301-2 / Vol", + " 138, MARCH 2016 Transactions of the ASME Downloaded From: http://mechanicaldesign.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jmdedb/934988/ on 02/21/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use system Sg Og : Xg; Yg; Zg to system Sw Ow : Xw; Yw; Zw\u00f0 \u00de is described as [9] Mwg \u00bc 0 0 1 aw cos p=2\u00fe P w cos p P w 0 0 cos p P w cos p=2 P w 0 0 0 0 0 1 2 664 3 775 (14) Then, the coordinates of arbitrary point A y; a;u\u00f0 \u00de on the tooth surface of gear (as shown in Fig. 3) are transformed to the coordinate system of the worm by the formula below: Aw xAw; yAw; zAw\u00f0 \u00de \u00bc MwgA xA; yA; zA\u00f0 \u00de (15) where Aw xAw; yAw; zAw\u00f0 \u00de are the coordinates of point A represented in coordinate system Sw Ow : Xw; Yw; Zw\u00f0 \u00de. A gear tooth surface is transformed to the coordinate system of worm by Eq. (15), as shown in Fig. 6. 3.4 Algorithm of Minimizing Optimization. It seems to be a difficult question to calculate the rotational angle of the worm that corresponds to a given angle of the gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000344_gt2015-43161-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000344_gt2015-43161-Figure1-1.png", + "caption": "Fig. 1 Schematic view of the herringbone-grooved journal bearing flexibly supported by straight spring wires", + "texts": [ + " In this paper, a flexibly supported herringbonegrooved aerodynamic journal bearing using straight spring wires made of stainless steel is proposed to provide a simple and reliable support system of a bearing bush. The threshold speed for the instability of aerodynamic bearings was assessed both experimentally and numerically. In addition, we investigated numerically the effect of dynamic stiffness and the damping coefficient of the straight spring wires on the stability of a rigid rotor supported by the proposed bearing. The proposed bearing flexibly supported by straight spring stainless steel wires is shown in Fig. 1. The proposed bearing consists of a rotor with herringbone grooves, a bearing bush, and straight spring wires to support the bearing bush. The bearing bush was supported by six straight spring wires that were assembled into a hexagon shape at the upper and lower positions, where the circumferential groove was formed on the outer surface. Both ends of the straight spring wires were fitted into narrow grooves machined on the sidewalls of the bearing housing, and spot-welded to the grooves. The groove depth was properly changed to assemble the wires into a hexagonal shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003798_icem49940.2020.9270715-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003798_icem49940.2020.9270715-Figure7-1.png", + "caption": "Fig. 7. The motor structure with different windings and rotor structures. For amorphous alloy materials, core loss curve was tested and compared with other materials. The comparison results are shown in Fig. 8.", + "texts": [ + " The effect of different winding structures on electromagnetic torque has been analyzed in section III. In fact, different winding structures also have a great effect on winding AC loss. Especially for high-speed SRMs, winding is supplied by high frequency current and winding harmonic content is larger than other motors, so the winding loss will also be serious. Moreover, different core materials and rotor structures will be analyzed, core loss and wind friction loss will be calculated respectively. The motor structure with different windings and rotor structures is shown in Fig. 7. For wind friction loss, the Reynolds number eR can be expressed as: e RtR \u03c9 \u03bc \u03c1 = (5) The motor angular velocity \u03c9 , air viscosity \u03bc , air density \u03c1 and temperature is constant. And the drag coefficient dC can be written as: ( )d d 1 2.04 1.768ln eR C C = + (6) The wind friction loss can be calculated by: 4 3 dfrictionloss salientP k C R L\u03c0 \u03c1 \u03c9= (7) For the Rotor 1, salientk need to be determined by CFD, and for Rotor 2, salientk need satisfy this N-S equation: 2 2 1 0 1 0 1 0 zvp z r p p r \u03bc \u03c1 \u03c1 \u03b8 \u03c1 \u2202\u2202\u2212 + = \u2202 \u2202 \u2202\u2212 = \u2202 \u2202\u2212 = \u2202 (8) Therefore, wind friction loss of two rotor structures can be calculated, the result is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001738_j.procir.2015.12.103-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001738_j.procir.2015.12.103-Figure5-1.png", + "caption": "Fig. 5. Presentation of gear set", + "texts": [ + " The results of the study show, on the one hand, that if the gear set show a topography scatter, it is necessary to consider the topography of all teeth, in order to receive a good agreement with the measurement. On the other hand, the results can be interpreted as an optimization potential for the optimization of the excitation behavior. By applying an intentional scatter of the topography a reduction of the amplitudes of the gear mesh orders can be achieved. 4. Design of bevel gear set with topography scatter Hence, the optimization is applied on an exemplary automotive gear set with a hypoid offset (Fig. 5). The gear set is milled in the single indexing process and ground afterwards. The TE is investigated in ZaKo3D and is simulated with 30 positions per pitch. The load-free TE of the ideal ease-off geometry is shown in Fig. 5 (bottom). For the calculation study, two gear variants are analyzed; the reference variant (Fig. 5) and one variant where the ideal geometry is superimposed with a micro geometry scatter. The topography deviations are stochastically distributed and shown in Fig. 6 right bottom. The values are derived from deviations that occur in lapping processes, but can also be manufactured in a grinding process. In order to analyze the effect of the topography it is necessary to simulate 506 pitches for one tooth hunt. The diagrams on the right side of Fig. 6 show that the amplitudes of the gear mesh order are decreased for the complete torque range" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure1-1.png", + "caption": "Fig. 1. The three investigated cooling system designs.", + "texts": [ + " The main contribution of this paper is the extension to the previously done models, including comparison of three different cooling systems, wave-cooling duct, spiral cooling duct, and a way to cool the end winding here denoted squiggly ducts due to the irregular shape of the cooling ducts. 3-D models with coupled FE electromagnetic and CFD models are used where losses and temperatures are transferred between models until a steady state is achieved. The motor models are previously compared with measurements in [3] and [4]. The investigation aim to show what design modifications that show possibilities to reduce hot spots and thermal fluctuations. Another goal is to find input to a lumped parameters model (LPN) [8]. Three standard cooling solutions, illustrated in Fig. 1, are investigated: a) Wave-cooling duct inside in the aluminum casing frame (already modelled in the complex 2-way coupled model [3] and used as a reference here). b) Spiral cooling duct inside the aluminum casing frame (the same as modelled with LPN in [1]). c) Squiggly cooling duct covering potted end-regions. The cooling duct is integrated in to the flange of the motor. The three cooling duct designs (wave, spiral and end region squiggly) have been studied and compared by using the following models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003804_msnmc50359.2020.9255521-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003804_msnmc50359.2020.9255521-Figure2-1.png", + "caption": "Fig. 2. Visualization of industrial robot work in the RobotStudio software environment.", + "texts": [ + " 98 2020 IEEE 6th International Conference on Methods and Systems of Navigation and Motion Control (MSNMC) emphasis for positioning of the object of manipulation (OM) of OM is placed, representing a square plate 100x100 mm thickness 10 mm with a fixer for installation of steel plates (for change of mass of OM) the accelerometer on the basis of IMU (MPU9255) sensor 10 DOF that OM mounted in a point of cent of masses and is connected to the Raspberry Pi 3B controller. The Raspberry Pi controller 3B analyzes the data received from the accelerometer and displays the data using a developed program with a discreteness of 2 ms. There were 2 types of studies: 1) study of transportation of objects along vertical trajectory; 2) study of transport of objects along vertical-horizontal trajectory. A robotic cell has been developed and programmed in the software environment RobotStudio [44] (Fig. 2). This allowed for studies with different movement parameters. For the first case of vertical trajectory transportation, a comparison of the two acceleration plots along the z axis (Fig. 3) for accelerometer rigidly fixed in gripping devices and fixed on object of manipulation. Transportation in the vertical direction was carried out using an object weighing 0.5 kg at a speed of 1 m/s and an acceleration of 22 m/s2. As can be seen from Fig. 3 acceleration acting on the OM repeats the acceleration profile of the gripping devices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure15-1.png", + "caption": "FIG. 15. Bending crease mode deformation sequence for a diametrically creased disk (N = 2) with \u03b2 = 24\u25e6, R = 30, r = 0.5, t = 0.1, \u03b10 = 17\u25e6, and n = 10. As the indentation is increased the ridge expands radially towards the edge of the disk.", + "texts": [ + " A vertex displacement, \u03b4, is imposed and the ridge radius, , obtained numerically by enforcing the symmetry boundary condition [Eq. (5)] using MATLAB [22]. The corresponding strain energy is computed using Eq. (31). The indentation depth is then incremented by a small amount, d\u03b4. Using the ridge radius obtained from the previous step as an initial guess, the radius corresponding to an indentation of \u03b4 + d\u03b4 is obtained using the same numerical scheme. This process is continued to compute the complete strain energy-indentation behavior. Figure 15 shows the resulting deformation sequence for a disk with a diametrical crease (N = 2). Under increasing indentation, the ridge radius increases linearly (see Fig. 16) until it reaches the outer edge, when the disk snaps to the inverted state. Using the strain energy-indentation results, the indentation force is calculated using Eq. (23). The change of energy and indentation force, as a function of indentation depth, are shown in Fig. 17. Under increasing indentation, the energy increases monotonically until the ridge reaches the outer edge of the disk" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003961_j.measurement.2020.108909-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003961_j.measurement.2020.108909-Figure5-1.png", + "caption": "Fig. 5. Path generation.", + "texts": [ + " For the sake of simplicity and also for comparison with the conventional DBB circular test according to [12], we will restrict the following discussion to the generation of a path with spherical curve, although the approach could also be applied to generate paths with varying measuring radius. Given the kinematic model of the hexapod, the approach comprises principally following four steps: 1. The sensitivity information is derived and the poses in the working space are sought, which are most sensitive to the kinematics of the machine concerned (Fig. 5(a)). 2. The poses in step 1 are reordered, so that a feasible and collisionfree path could result (Fig. 5(b)). Criteria, that should be taken into account, include the curve length, the kinematic feasibility and kinetic constraints. 3. Equidistant support points for CNC control are generated along a smooth curve crossing the reordered poses in step 2 (Fig. 5(c)). 4. The orientation for support points in step 3 is interpolated and optimized, so that the resulting path is feasible, free of collision and as smooth as possible in joint space. We measured the generated path (Fig. 6) and sampled the values with a nominal rate of 250 Hz in a dedicated data capture device, which we will introduce in the next section. The captured DBB values (Fig. 7, bottom) exceed the measuring range of the commercial DBB (e.g. Renishaw QC20-W [10]), which confirms our claim at the beginning of current section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure6-1.png", + "caption": "Figure 6. A cross-sectional diagram and side view of the fabricated initial gear prototype with conical passive rollers", + "texts": [ + " In the following chapters, we introduce the specifications of two real prototypes of gears with passive rollers. This time, we fabricated a gear with conical passive rollers and one with flat passive rollers that can be manufactured economically and with simple mechanisms. To test the real features of the gear with passive rollers, we initially fabricated a gear with conical passive rollers because of the considerable scope for miniaturization. A cross-sectional diagram of this designed gear with conical rollers is shown in Fig. 6, while a cross-sectional diagram of one conical passive roller unit and an exploded view of its components are shown in Fig. 7. To achieve smooth sliding motion with the omnidirectional driving gear, the gear structure with conical passive rollers requires multiple layers with some offset, since the omnidirectional driving gear has an intermittent structure of multiple teeth as a rack gear. Accordingly, when in actual use, the gear with a conical passive roller should be designed with at least three layers [4]. However, for the first trial, we made the prototype with only one layer of conical passive rollers, as shown in Fig. 6 and 8, to determine its basic feasibility and mechanical advantage. 200,32 )2(cossin )1(sincos A Ay Ax The outline of the conical passive roller is manufactured by a CNC milling machine according to theoretical equations of involute curve as shown in equations (1) and (2). These equations show the description of x and y coordinates on the red involute curve in Fig. 7. The origin of the x-y coordinate is placed on the starting point of the involute curve of the conical passive roller at its bottom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001843_jahs.59.022006-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001843_jahs.59.022006-Figure3-1.png", + "caption": "Fig. 3. Inboard bell crank kinematic diagram.", + "texts": [ + " While the configuration used here has two bell cranks and two sets of linkages, which could all be modified in various ways to change the kinematics, the design used in this study employed linkages with equal moment arms on both ends, such that they directly transmit the torque and deflections of the inboard bell crank without modification. Thus, the PAM flap system kinematics are solely determined by the mounting of the PAMs to the inboard bell crank. The primary parameters of interest for these kinematics are the changes in length and effective moment arms for both the active and passive PAMs as a function of the flap deflection angle \u03b4. The solid black lines in Fig. 3 show the simplified geometry of the bell crank at rest, with the point PM being the mounting point for the active PAM, point PN being the mounting point for the passive PAM, and point PO being the rotation point of the bell crank. The dashed lines represent this same bell crank rotated through a positive angle \u03b4. The effective moment arm of the active PAM, xA, will be the chordwise distance between PM and PO when the bell crank is rotated through the angle \u03b4: xA = xoff cos (\u03b4) + yoff sin (\u03b4) (13) Similarly, the effective moment arm of the passive PAM xP , which is mounted at PN , is xP = xoff cos (\u03b4) \u2212 yoff sin (\u03b4) (14) The change in length of the two PAMs is the difference between the spanwise projection of the PAM-mounting points for the bell crank at rest and when deflected by the angle \u03b4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003711_j.eml.2020.101076-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003711_j.eml.2020.101076-Figure2-1.png", + "caption": "Fig. 2. Four motion states for one sub-process of passive walking: (1) left singleeg support phase; (2) double-leg support phase; (3) right single-leg support hase; (4) double-leg support phase.", + "texts": [ + " The cross section of the leg is rectangular with width w and length d, and the sectional area is A = w \u00d7 d. The leg has a cylinder hole with radius R, which is used to assemble the pin. The material parameters of the leg are Young\u2019s modulus El, Poisson ratio \u03bdl and density \u03c1l. The slope is simplified to six footplates with length a, width b and height . The material parameters of the slope are Young\u2019s modulus Es, oisson ratio \u03bds and density \u03c1s. The COF between the foot and the slope is \u00b5l-s and the COF between the leg and the pin is \u00b5l-p. As shown in Fig. 2, the whole walking is specified as a process that consists of multiple repeated sub-processes. Each subprocess is further divided into four phases, which are (1) left single-leg support phase, i.e. the left leg is in contacting with the slope, the right leg is swinging; (2) double-leg support phase, i.e. the left leg and the right leg are both in contacting with the slope; (3) right single-leg support phase, i.e. the right leg is in contacting with the slope, left leg is swinging; and (4) doubleleg support phase, i", + " The COF between the leg and the pin is ignored. Meanwhile, the slope is modeled as a rigid ground. The connection at the hip is modeled as a plane joint and the mass of the pin is modeled as a concentrated mass attached on the pin. The mass of the pin and the leg are mp = \u03c1p \u00d7\u03c0 \u00d7 r2 \u00d7h and m = \u03c1l \u00d7 l\u00d7A, respectively. The distance from the center of leg mass to the foot is ld and the 2 moment of inertia of the leg is J = ml /12 as ld = l/2. l p a o s s i 2 ( I g w o p o x n a i l s c t w s t a M w p H ( w f Since the phases 1 and 3 in Fig. 2 have the same contact state nd the phases 2 and 4 have the same contact state, the walking f the CRBM is further classified into two phases, which are (1) ingle-leg support phase, i.e. support leg is in contacting with the lope, swing leg is swinging; and (2) double-leg support phase, .e. two legs are both in contacting with the slope. .1.1. Dynamic equations of CRBM 1) Dynamic equations during single-leg support phase Fig. 3 shows the CRBM is during the single-leg support phase. t is assumed that there is no slip at the contact point", + " Then based on the conservation of angular momentum, the motion states of the legs after the impact are calculated by the following equation, Q+ (2\u03b2) q\u0307+ 2 = Q\u2212 (2\u03b2) q\u0307\u2212 2 (5) where 2\u03b2 is the angle between two legs at the impact moment and Q is the generalized moment of inertia. 2 T m s m b p \u0398 E 2 w m e t i w n 2 e t t s m a M w m u t f m u w t m \u03c3 w e 2 g T s o n n t s t h{ w i m o a l S t r w o t c g c e \u03b4 w n a o p T f w s w o t r T n .1.2. Initial states of single-leg support phase for stable walking The walking is a periodic motion problem as shown in Fig. 2. hrough equation (2), the system motion states at the moveent moment (i.e. single-leg support phase) at the nth single-leg upport phase can be calculated. Then substituting these system otion states into Eqs. (4) and (5), the initial states of the (n+1)th single-leg support phase can be obtained. Hence, the initial states of the (n + 1)th single-leg support phase \u0398n+1 can be calculated y a function of the initial states of the nth single-leg support hase \u03d5(\u0398n) as n+1 = \u03d5 (\u0398n) (6) where \u0398n = [\u03b8 (n) 1 \u03b8 (n) 2 \u03b8\u0307 (n) 1 \u03b8\u0307 (n) 2 ] T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000835_jjap.53.01ae03-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000835_jjap.53.01ae03-Figure1-1.png", + "caption": "Fig. 1. (Color online) Basic schematic of electrospinning.", + "texts": [ + "23\u201325) It is easy to produce nanofibers with a variety of properties by electrospinning, with the physical properties of a very small diameter and large surface area per unit mass, orientated nanofibers produced by electrospinning are expected to affect the orientation of LCs. In this study, we report on a twisted nematic LC cell with an alignment layer of orientated nanofibers instead of rubbed polyimide. The optical characteristics of this cell are measured and compared with those of a conventional rubbed polyimide cell. The polymer used in this experiment was poly(vinyl alcohol) (PVA) with a polymerization degree of 500. PVA was dissolved in water to a concentration of 25wt%. A basic setup for the electrospinning process is shown in Fig. 1. Details of the parameters of the electrospinning setup for producing nanofibers are given in Table I. Polymer jets are easily released by applying a high voltage (typically from 1 to 30 kV) between the collector and the nozzle. When the applied electric field increases and reaches a certain value, the electric force overcomes the surface tension of the polymer solution and forces the liquid jet to stretch out. The liquid jet continuously elongates and the solvent evaporates during the process. Finally, a random formation of nonwoven fibers is observed on the stationary collector,26) as shown in Fig. 2(a). To obtain orientated nanofibers by electrospinning, we used a drum collector (Normal Drum in Fig. 1) and embanked a drum collector (Improved Drum in Fig. 1). During the electrospinning process, the drum rotates at a very high speed so that nanofibers can be taken up on the surface and wound around it. Finally, we expected nanofibers to be oriented along the rotation direction. By mounting an indium\u2013tinoxide (ITO) glass substrate on the surface of the collector during the electrospinning process, we also obtained a thin layer of nanofibers on the ITO surface. Japanese Journal of Applied Physics 53, 01AE03 (2014) http://dx.doi.org/10.7567/JJAP.53.01AE03 REGULAR PAPER 01AE03-1 \u00a9 2014 The Japan Society of Applied Physics Aligned nanofibers collected from the normal drum are shown in Fig", + " The cells were then kept between two parallel polarizers, the (a) (b) (c) 300 \u00b5m Rotation direction Power spectraMicrographs Fig. 2. Micrographs of nanofibers prepared by electrospinning and their power spectra: (a) random nanofibers, (b) nanofibers from normal drum at rotation speed of 2000 rpm, and (c) nanofibers from improved drum at rotation speed of 1000 rpm. Fig. 3. (Color online) Calculated distributions of electric field on the surfaces of normal and banked collectors. Table I. Parameters for producing nanofibers for setup shown in Fig. 1. Parameter Random nanofibers Nanofibers from normal drum Nanofibers from embanked drum PVA500 concentration (wt%) 25 25 25 Nozzle\u2013collector surface distance (mm) 100 100 75 Applied voltage (kV) 25 25 25 Feed rate of PVA (ml/h) 0.2 0.2 0.2 Rotation speed (rpm) 0 2000 1000 Time of process (s) 60 300 300 01AE03-2 \u00a9 2014 The Japan Society of Applied Physics rubbing direction of one glass plate was parallel to the polarizers. Rectangular voltage at a frequency of 1 kHz supplied from a function generator (Tektronix AFG 3022) was applied to the cells" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003365_lra.2020.3010444-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003365_lra.2020.3010444-Figure2-1.png", + "caption": "Fig. 2. CAD models of the nine 3 d printed object shapes.", + "texts": [ + " The output of this stage is a 1 d channel image with a binary probability at each pixel location indicating whether a grasp at that location is likely to succeed or not. Objects can have many different intrinsic properties. In this work, we focus on how an object\u2019s mass, and mass distribution affects the performance of grasping with a suction cup gripper. To account for real-world phenomenon (e.g. the flexibility of a gripper\u2019s suction pad), we collect data in the real-world using a 6-DoF Fanuc 200iC arm, along with nine custom 3 d printed object shapes as shown in Fig. 2. Many real-world objects do not readily allow access to their intrinsic properties, nor do they easily allow for these properties to be re-configured in any consistent way. We originally thought to construct different shapes through e.g. lego-like parts, but found it difficult to build objects that were able to support additional weight and consistently hold their shape after being dropped by the gripper. By 3 d printing the objects, we were able to standardize these parameters. Our objects are simple, planar objects, and share some similarity to those used in [21]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure8-1.png", + "caption": "Fig. 8 The contact situation when peripheral cutting edge takes part in machining", + "texts": [ + " (10\u201313), we can obtain the mathematical expressions of boundary curve A-B-C in Fig. 7, and cutting depth of any IEi is just the z coordinates of curve A-B-C. Line A-B can be expressed as, z zA y yA \u00bc tan \u03b1 zA \u00bc 0; yA \u00bc e R ap tan \u03b1 2 ; y\u2208 yA; yB\u00bd Arc B-C can be expressed as, z zO1\u00f0 \u00de2 \u00fe y\u2010yO1\u00f0 \u00de2 \u00bc R2 zO1 \u00bc \u2212 R ap ; yO1 \u00bc e; y\u2208 yB; yC\u00bd So expressions of cutting depth z is, z \u00bc y e\u00fe R ap tan \u03b1 2 h i tan \u03b1 y\u2208 yA; yB\u00bd ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 y e\u00f0 \u00de2 q \u00fe ap\u2212R y\u2208 yB; yc\u00bd 8<: \u00f014\u00de Figure 8 shows the case of peripheral cutting edge E-C taking part in machining. Shaded region ADBEC is the cutting region on plane YOZ. For any IEi, whose Y coordinate is located between points B and C, the calculation of its cutting depth zB~D becomes complex. To simplify the calculation, we can move point D toD\u2032, as shown in Fig. 8b, making points B and D\u2032 share the same Y coordinate. Then the cutting region changes from ADBEC to AD\u2032BEC. Similarly, we can get expression of cutting depth z with region AD\u2032BEC by obtaining the coordinates of key points A, B, C, D\u2032, E and workpiece center O1, yA \u00bc e R ap tan \u03b1 2 zA \u00bc 0 ( \u00f015\u00de yB \u00bc Rc e\u2212 R ap tan \u03b1 2 h i cos \u03b1 \u00fe e R ap tan \u03b1 2 \u2212 ap R\u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 Rc e\u00f0 \u00de2 q sin \u03b1 zB \u00bc Rc e R ap tan \u03b1 2 h i sin\u03b1 \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 Rc e\u00f0 \u00de2 q R\u00fe ap cos \u03b1 8>>>>>>><>>>>>>>: \u00f016\u00de yC \u00bc Rc zC \u00bc ap R ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi R2 Rc e\u00f0 \u00de2 q 8<: \u00f017\u00de yD0 \u00bc yB zD0 \u00bc zD \u00bc RC e R ap tan \u03b1 2 h i sin \u03b1 ( \u00f018\u00de yE \u00bc Rc zE \u00bc 0 \u00f019\u00de yO1 \u00bc e zO1 \u00bc R ap \u00f020\u00de According to Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001131_icrom.2013.6510116-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001131_icrom.2013.6510116-Figure11-1.png", + "caption": "Figure 11 - Robot motion time history for CASE III", + "texts": [], + "surrounding_texts": [ + "The stabilization of a two Wheel Mobile Manipulator (WMM) was proposed using a reaction wheel. A virtual inverted pendulum was modelled by a 4-DOF manipulator, and it was controlled to achieve the robot balancing. The CoG position of the robot was identified by observing the reaction wheel position. By angular velocity and acceleration of the reaction wheel, the set point of PID controller was changed to the <11 position using the supervising control. This control algorithm was performed to achieve the CoG on the desired location and make the robot stable. To decrease the control calculation, this method was never used the calculation of the CoG identification. However, the supervising control method is used. Finally, simulation results proved the effectiveness of reaction wheel employing. This simulation shown that robot could start in any position and made itself stable. 4ur-------------,I---------------,I--------------\ufffdI I 30 k- - - - - - - - \ufffd - - - - - - - - - - - - \ufffd - ----------- .. ... \u00b7--\u00b7\u00b7\u00b7-\u00b7\ufffd\u00b7 ..... \u00b7r\u00b7-\u00b7-\u00b7 I 20 : - - - - - - - - - - - \ufffd - - - - - - - - - - - - \ufffd - ----------- 4U,-------,-------\ufffd------_,--------,_--\ufffd==\ufffd======\ufffd 30 ______ :______ \ufffd ______ \ufffd ______ :____ I =:- \ufffd:\ufffd\ufffd:;\ufffd\ufffdIe l 20 , , , -------------------------, , , 10 . - - - - - _1 - _____ -.J ______ l ______ 1 _______ 1 _____ _ E o - - - - - -1- - - - - - ---I - - - - - - \ufffd - - - - - - \ufffd - - - - - -1------ \ufffd -10 - - - - - -1- - - - - - ---t - - - - - - t- - - - - - - r- - - - - - -1------ -20 - - - - - -1- - - - - - I - - - - - - T - - - - - - r - - - - - -1- ----- I I I I I -30 \\- -----:- - - - - - \ufffd - - - - - - \ufffd - ----- :-------:------ '\" ____ ,I\"r ,'- ' , , -40 ---\ufffd------\ufffd \ufffd------:r--\"\"'''''''''''''''l -500\ufffd------,\ufffd0\ufffd----\ufffd2\ufffd0-------3\ufffd0\ufffd----\ufffd4\ufffd0-------=50\ufffd----\ufffd6' Time(s} Suppression of Two-Wheel Mobile Manipulator Using Resonance Ratio-Control-Based Null-Space Control\", iEEE Transactions on industrial Electronics, 57(12), pp. 4137-4146, (2010) [2] Nguyen, H. G., Morrell, 1., Mullens, K., Bunneister, A., Miles, S., Farrington, N., Thomas, K., and Gragee, D. W. \"Segway Robotic Mobility Platform\", Proc. SPiE Mobile Robots, Philadelphia, (2004) [3] Huang, Ch. and Fu, Li., \"Passivity based control of the double inverted pendulum driven by a linear induction motor\", Proc of iEEE Conference on control Applications, 2, pp. 797-802 (2003) [4] Grasser, F., D' Arrigo, A., Colombi, S., and Rufer, A. c., \"JOE: A Mobile, Inverted Pendulum\", iEEE Trans. on industrial Electronics, 49(1), pp. 107-114, (2002) [5] Niki, H. and Toshiyuki, M., \"An Approach to Self Stabilization of Bicycle Motion by Handle Controller\", Proc of the first Asia international Symposium on Mechatronics, (2004) [6] Rubi, J., Rubio, A., and Avello, A., \"Swing-up control problem for a self-erecting double inverted pendulum\", iEEE Proc. Control theory Application, 149(2), pp.169-175 (2002) [7] Liu, D., Gue, W., Yi, 1., and Zhao, D., \"Double-Pendulum-Type Overhead Crane Dynamics and Its Adaptive Sliding Mode Fussy Control\", Proc. Of the Third into Conf. on Machine Learning and Cybernetics, pp.26-29, (2004) [8] Pathak, K., Franch, J., and Agrawal, S. K., \"Velocity and position control of a wheeled inverted pendulum by partial feedback linearization\", iEEE Trans. RobotiCS, 21(3), pp. 505-513, (2005) [9] Zhang, M. and Tran, T., \"Hybrid Control of the Pendubot\", iEEE/ASME Trans. Mechatronics, 7(1), pp. 79-86 (2002) [10] Ibanez, C. A., Frias, O.G., and Castanon, M. S., \"Lyapanov-based controller for the inverted pendulum cart system\", Nonlinear Dynamics, 40(4), pp. 367-374, (2005) [11] Ge, S.S., Hang, C. c., and Zhang, T., \"A direct adaptive controller for dynamic system with a class of nonlinear parameterizations\", Automatica, 35, pp. 741-747 (1999) [12] Brooks, R., Aryanada, L., Edsinger, A., Fitzpatrick, P., Kemp, C. c., Reilly, 0., Torres-jara, E., Varshavskaya, P., and Weber, J., \"Sensing and manipulating built-for human environments\", int. J. of Humanoid RobotiCS, 1(1), pp. 1-28 (2004) [13] Miyashitaa, T. and Ishiguroa, H., \"Human-linke natural behavior generation based on involuntary motions for humanoid robots\", Robotics Autonomous Systems, 48, pp. 203-212 (2004) [14] Gans, N. R., and Hutchinson, S. A., \"Visual servo velocity and pose control of a wheeled inverted pendulum through partial-feedback linearization\", Proc. of iEEE/RSJ international Conference on intelligent Robots and Systems, pp. 3823-3828 (2006) [15] Yamamoto, Y. and Yun, x., \"Coordinating Locomotion and Manipulation of Mobile Manipulator\", Proc. of the 31st iEEE Conference on Decision and control, pp. 2643-2648 (1992) [16] Yamamoto, Y. and Yun, x., \"Control of Mobile Manipulators Following a Moving Surface\", Proc. of iEEE international Conference on robotics and Automation, 3, pp. 1-6 (1993) [17] Seraji, H., \"An On-line Approach to Coordinated Mobility and Manipulation\", Proc. of iEEE international Conference on Robotics and Automation, 1, pp. 28-35 (1993) [18] Acar, C. and Murakami, T., \"Multi-Task Control for Dynamically Balanced Two-Wheeled Mobile Manipulator through Task-Priority\", iEEE international Symposium on industrial ElectroniCS, pp. 2195- 2200 (2011) [19] Sasak, K. and Murakami, T., \"Pushing operation by two-wheel inverted mobile manipulator\", iEEE international Workshop on Advance Motion Control, pp. 33-37, (2008) [20] Alipour, Kh. And Moosavian, S. Ali. A., \"How to Ensure Stable Motion of Suspended Wheeled Mobile Robots\", Journal of industrial Robot, 38(2), pp. 139-152 (20 I I) 50,--------------.--------------,--------------. , , o ----------- t- ----------- t- ----------- -50 - - - - - - - - - - - L ___________ L - ---------- \ufffd -100 :3- , , - - - - - - - - - - -1 -- - - - - - - - - - T - ---------- -150 <>: - - - - - - - - - -1- - - - - - - - - - - - +- - ---------- \ufffd '0 -200 -250 r ' , , - - - - - - - - - -1 -- - - - - - - - - -1 - ---------- -300L--------------L--------------\ufffd------------\ufffd o 10 15 lime (s) Figure 15 - Position of the reaction wheel 100 .--------------.--------------.--------------, o -----------\ufffd-----------\ufffd----------- ___________ L ___________ L __________ _ -; -200 - - - - - - - - - - - f- - - - - - - - - - - - f- ----------- c;, c: ..: -300 -400 , , -------1 -----------1 ----------- - - - - - - - ----- -\ufffd- - --\ufffd - - - - - -----\"\"\"'- -500 '--____________ L-____________ -'-____________ --' o 5 10 15 lime (5) -50 0 \ufffd- - - ,\ufffd0- - - \ufffd20\ufffd- - \ufffd 3\ufffd0- - - \ufffd 40\ufffd- - \ufffd5\ufffd0 - - - \ufffd60 Time(s) Figure 17 - Position of the reaction wheel" + ] + }, + { + "image_filename": "designv11_34_0000803_iros.2013.6696525-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000803_iros.2013.6696525-Figure1-1.png", + "caption": "Figure 1. Link frame of the KR270 robot and picture of robot", + "texts": [ + " An unbiased estimation of the standard deviation \u02c6W is [14]: \u02c6 /W cs r c (19) The covariance matrix of the estimation error is approximated by [14]: 1: 1 1: 1 122 \u02c6 \u02c6 1: 1 2 \u02c6 \u02c6\u02c6\u02c6 1 c c T W c tot totC W W (20) with 1: 1\u02c6 c the vector containing the 1c first coefficients of *\u02c6 tot and 1: 1 \u02c6 ctotW a matrix composed of the 1c first columns of \u02c6 totW . Finally, 2 \u02c6 \u02c6 \u02c6 ( , ) i C i i is the ith diagonal coefficient of \u02c6 \u02c6C and the relative standard deviation \u02c6% ri is given by: \u02c6 \u02c6 \u02c6% 100 ri i i , for \u02c6i \u2260 0. \u02c6 \u02c6 \u02c6 \u02c6% 100 for 0 ri i i i (21) V. MODELING THE KUKA KR270 AND ITS SPRING BALANCER A. Modeling of the Kuka KR270 The Kuka KR270 (see fig. 1) robot has a serial structure with 6 rotational joints. Its kinematics is defined using the MDH notation described in section II-A. The geometric parameters defining the robot frames are given in table I. All mechanical variables are given in SI unit in joint side. The robot is characterized by a kinematic coupling effect between the joint 4,5 and 6: 4 5 6 4 4 54 5 5 64 65 6 6 0 0 0 m m m q K q q C K q C C K qq (22) With mjq is the motor velocity of joint j and jq is the joint velocity of joint j", + " The values 54 64 65( , , )C C C are very low (factor 100) compared to the values 4 5 6( , , )K K K , therefore the kinematic coupling effect is not considered in the dynamic modeling. The support of payload and payload are shown on fig. 2 B. Modeling of the spring balancer This robot has a spring balancer attached to the second joint. It is composed by a spring and induces a torque on the joint 2. This system compensates the gravity joint torques caused by the weight of robot links 2 to 6. The location of the system on the robots is shown in figure 1. The torque induced must be taken into account in the dynamic model of the robot to identify all parameters accurately. The kinematics of the spring balancer uses the MDH notation in order to compute the compensation torque applied on joint 2. The MDH parameters defining the spring balancer frames are given in table II and its kinematics are detailed in figure 3. 22 22 22 1 1 2 1 2 cos cos sin 90 sb F F q d d F dq (24) With: 23 22 212 90q qq q q (25) Consequently, the values 221q q , 4 2rl q and 223q q must be expressed function of 2q " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.3-1.png", + "caption": "Figure 10.3 A planar parallel manipulator formed by two planar serial manipulators.", + "texts": [ + " s1 = s01 = OD1, s2 = s02 = OD2, s3 = s03 = OD3 s4 = s11\u22158 = B1A1, s5 = s12\u22159 = B2A2, s6 = s13\u221510 = B3A3 Then, the secondary variables happen to be the passive joint variables associated with the 12 revolute joints. They are the angles defined similarly as in Example 10.1. The four IKLs of this parallel manipulator can be taken as follows: IKL-1 = D1D2C2C1D1, IKL-2 = D2D3C3C2D2 IKL-3 = B1B2A2A1B1, IKL-4 = B2B3A3A2B2 Example 10.3 Planar Parallel Manipulator Formed by Two Planar Serial Manipulators The planar parallel manipulator shown in Figure 10.3 is formed as a combination of two coplanar serial manipulators that are collaborating in order to manipulate a rigid bar that they tightly grip. Owing to the tight grip, the rigid bar and the two grippers may be assumed to be integrated into a single link, which becomes the moving platform of the manipulator. Thus, nm = 5 and j1 = 6. Therefore, according to Eqs. (10.6)\u2013(10.9), \ud835\udf07 = 3, nikl = 1, npv = 3, nsv = 3 Note that this parallel manipulator has only one kinematic loop, which is O1O2A2B2 B1A1O1, and its mobility is \ud835\udf07 = 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002229_acc.2016.7525039-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002229_acc.2016.7525039-Figure7-1.png", + "caption": "Figure 7. Logarithmic norm of the complementary sensitivity function vs. logarithmic dimensionless frequency, \u03c4\u0304\u03c9, for different values of \u03b5. \u03b3 = \u22120.3 and \u03bb \u03c4\u0304 = 1.5. The surfaces are plotted for three different levels of gain relative error, \u03b4 = 0,\u00b10.3. Complementary sensitivity function now exhibits a better performance with zero steady-state error at low frequencies (|\u03b7(j\u03c9)| \u2248 1) and smaller magnitude at higher frequencies.", + "texts": [], + "surrounding_texts": [ + "To design the controller, we will first introduce a lowpass filter into the controller dynamics in order to roll off the \u03b7(s) at high frequencies under a load of unit step. It also helps to prevent large or rapid control actions. Let consider the filter as f(s) = 1 (\u03bbs+ 1) n . (7) To avoid excessive differential actions we will consider a first order filter, i.e., n = 1. Choosing various stable GC(s) parameterizes all stabilizing controllers. This Youla parameterization is valid for unstable systems as well [31]. Hence, the controller is designed as a cascade of stable and causal inversion of the delay-free nominal system and a filter given by (7). The \u03bb-tuned controller is then formed as GC(s) = T\u0304 s+ 1 K\u0304 \u00b7 1 \u03bbs+ 1 , (8) with \u03bb being the time-constant of the low-pass filter. A systematic framework to tune the parameter \u03bb with the permissible values is provided by the following theorem. Theorem 1. Consider a family of systems in \u2126 given by (1). Define \u03b4 , K \u2212 K\u0304 K\u0304 , \u03b5 , \u03c4 \u2212 \u03c4\u0304 \u03c4\u0304 , and \u03b3 , T \u2212 T\u0304 T\u0304 as uncertainty in the system gain, delay, and time-constant, respectively. Hence, the prescribed bounds given by (2) can be rephrased as |\u03b4| \u2264 p, (9a) |\u03b5| \u2264 q, (9b) |\u03b3| \u2264 r. (9c) Then the controller (8) provides (3) with robust stability if the following inequality holds { \u03b42 + 4(\u03b4 + 1) [ sin \u03b5\u03c9\u03c4\u0304 2 + (\u03b3 + 1)(\u03c9\u03c4\u0304)2 ( T\u0304 \u03c4\u0304 )2 sin \u03b5\u03c9\u03c4\u0304 2 + \u03b3(\u03c9\u03c4\u0304) ( T\u0304 \u03c4\u0304 ) cos \u03b5\u03c9\u03c4\u0304 2 ] sin \u03b5\u03c9\u03c4\u0304 2 + (\u03c9\u03c4\u0304)2 [ T\u0304 \u03c4\u0304 (\u03b4 \u2212 \u03b3) ]2} 1 2 \u00d7 { 1 + [ \u03c9\u03c4\u0304(\u03b3 + 1) T\u0304 \u03c4\u0304 ]2}\u2212 1 2 < [(\u03bb \u03c4\u0304 \u03c9\u03c4\u0304 )2 + 1 ] 1 2 . (10) Proof: For an uncertain system satisfying the multiplicative uncertainty condition given by (1), the small-gain theorem \u2016\u03b7\u0304\u2206\u2016\u221e < 1, (11) ensures the internal stability. Using appropriate substitutions, Eq. (11) can be expressed as\u2223\u2223\u2223GC(j\u03c9)G\u0304P (j\u03c9) GP (j\u03c9)\u2212 G\u0304P (j\u03c9) G\u0304P (j\u03c9) \u2223\u2223\u2223 = \u2223\u2223[GP (j\u03c9)\u2212 G\u0304P (j\u03c9) ] GC(j\u03c9) \u2223\u2223 < 1. (12) Substituting (1) and (8) into (12) gives\u2223\u2223\u2223\u2223(Ke\u2212j\u03c4\u03c91 + jT\u03c9 \u2212 K\u0304e\u2212j\u03c4\u0304\u03c9 1 + jT\u0304\u03c9 ) \u00b7 jT\u0304\u03c9 + 1 K\u0304 \u00b7 1 j\u03bb\u03c9 + 1 \u2223\u2223\u2223\u2223 = \u2223\u2223\u2223\u2223K(1 + jT\u0304\u03c9)e\u2212j\u03c4\u03c9 \u2212 K\u0304(1 + jT\u03c9)e\u2212j\u03c4\u0304\u03c9 K\u0304(1 + jT\u03c9) \u00b7 1 j\u03bb\u03c9 + 1 \u2223\u2223\u2223\u2223 < 1. (13) After carrying out some mathematical manipulations, the above inequality becomes{(K \u2212 K\u0304 K\u0304 )2 + 4 K K\u0304 sin2 \u03c9(\u03c4 \u2212 \u03c4\u0304) 2 + (\u03c9T\u0304 )2 [(K K\u0304 \u2212 T T\u0304 )2 + 4 K K\u0304 T T\u0304 sin2 \u03c9(\u03c4 \u2212 \u03c4\u0304) 2 ] + 2\u03c9 K K\u0304 (T \u2212 T\u0304 ) sin\u03c9(\u03c4 \u2212 \u03c4\u0304) } 1 2 \u00d7 {[ 1 + (\u03c9T )2 ][ 1 + (\u03bb\u03c9)2 ]}\u2212 1 2 < 1. (14) Substituting uncertainty definitions for the parameters in (14) yields the inequality (10) and completes the proof. Inequality (10) presents a direct rule to choose the tuning parameter, \u03bb, in order to meet the robust stability conditions as a function of nominal delay. For timevarying nominal delays such an inequality can yield a less conservative result. This inequality involves dimensionless ratios T\u0304 \u03c4\u0304 (nominal time-constant over nominal delay), \u03bb \u03c4\u0304 (tuning parameter or filter time-constant to nominal delay), and independent variable \u03c9\u03c4\u0304 (dimensionless frequency) to describe the robust stability requirement. The nominal values of the parameters are presumably known to the designer. Moreover, when there is only uncertainty on the delay, i.e., \u03b5 6= 0, \u03b3 = 0, and \u03b4 = 0, one can obtain the following inequality: 1 \u03c9\u03c4\u0304 \u221a 4 sin2 \u03b5\u03c9\u03c4\u0304 2 \u2212 1 < \u03bb \u03c4\u0304 . (15) Further investigation on the performance of the proposed IMC at various operating conditions is made possible through Eq. (10)." + ] + }, + { + "image_filename": "designv11_34_0003137_med48518.2020.9183255-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003137_med48518.2020.9183255-Figure2-1.png", + "caption": "Fig. 2. Kinematic model of the vehicle.", + "texts": [ + " Paper structure. Section II introduces the kinematic model and the problem statement. Section III synthesizes the controller design and analyses the stability of the closed-loop system. Section IV offers simulation and experimental results using the AutoNOMO vehicle. Finally, Section V provides some concluding remarks. The kinematic model of the vehicle, taken from [12], is given as x\u0307 = v1 cos(\u03b8), y\u0307 = v1 sin(\u03b8), \u03b8\u0307 = v1 ` tan(\u03d5), \u03d5\u0307 = v2, (1) where x(t) \u2208 R, y(t) \u2208 R are the Cartesian coordinates (see Fig. 2), \u03b8(t) \u2208 R is orientation angle with respect to the X-axis and \u03d5(t) \u2208 R is the steering angle, v1(t) \u2208 R and v2(t) \u2208 R are the control inputs, v1 corresponds with the linear velocity and v2 is the steering velocity, ` is the distance between the front and rear wheels. In this work it is assumed that all the variables are measured. The vehicle has a restriction in the steering angle rotation due to the mechanical constraints, i.e., |\u03d5| \u2264 \u03d5\u0304 with \u03d5\u0304 \u2248 0.66[rad]. Moreover, the vehicle presents bounded linear and steering velocities, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000963_nems.2014.6908895-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000963_nems.2014.6908895-Figure2-1.png", + "caption": "Figure 2. (a) A photograph of the microfluidic chip, which was composed of several micro-components. Note that red region and blue region are the liquid layer and the air layer, respectively. (b) An exploded view of the microfluidic chip and a schematic diagram of the all bio-sample partition on the microfluidic chip.", + "texts": [ + ", Germany) and via high-performance liquid chromatography (HPLC, Bio-Rad Variant II Turbo), respectively. 978-1-4799-4726-3/14/$31.00 \u00a9 2014 IEEE 647 B. The design of microfluidic chip In order to perform the sandwich assay to measure Hb and HbA1c from blood samples (Figure 1), we developed a trilayer microfluidic chip which consisted of two polydimethlysiloxane (PDMS, Sylgard 184A/B, Dow Corning Corp., USA) layers including a liquid channel layer (thickfilm) and a pneumatic layer (thin-film) layer on one glass plate (G-Tech Optoelectronics Corp., Taiwan). As shown in Figure 2, the major micro-components such as normally-closed valves, a transport unit, a closed chamber, and five open chambers were integrated into this device. The dimensions of the chip were measured to be 30 \u00d7 28 mm. The microfluidic chip employed a vacuum pump and an air compressor controlled by an electromagnetic valve (EMV) to drive and mix the sample and reagents fluids in the incubation chamber. The pumping characteristic of the microfluidic control module is shown in Figure 3. The pumping rate of the micropump is proportional to the applied suction pressure below 66" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002293_jae-162048-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002293_jae-162048-Figure3-1.png", + "caption": "Fig. 3. Operating principle.", + "texts": [ + " The structure of the HTLA without axially stacking permanent magnets has advantages against conventional cylindrical linear actuators. In the proposed structure composed of the helical teethed yokes with interior permanent magnets, the number of teeth can be increased without increasing the number of permanent magnets. When the 3-phase coils are excited, the mover rotates the same way as a rotary motor. The rotation changes the relative position between the teeth of the mover and stator as shown in Fig. 3(a). Thereby the phase difference is changed and magnetic attractive force is induced to the mover as shown in Fig. 3(b). In the vector control, the directly controlled variables are the amplitude and phase of the current. The relative position between the stator and mover is related to the thrust force generation. The phase difference shown in Fig. 3(b) is defined as follows: \u03c6 = NT ( \u03b8 \u2212 2\u03c0 L z ) (1) where NT is the number of helical threads, L is the distance per revolution on helical teeth, z is the position and \u03b8 is the rotation angle. The static characteristics were calculated by 3-D FEM. The major specifications of HTLA are shown in Table 1. We calculated the thrust force and torque when the rotation angle was changed from the \u221290 deg. to 90 deg. in steps of 5 deg. While the linear position is fixed, the phase difference was changed. The input average q-axis current Iq under the vector control was 0 or 1 A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003405_rpj-07-2019-0200-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003405_rpj-07-2019-0200-Figure1-1.png", + "caption": "Figure 1 LPBF testing geometry with dimensions and print strategy", + "texts": [ + " However, the porosity introduced by this reduced infill can be distinguished from the unintentional process-related porosity as shown in Figure 3. The LBPF samples were printed on an ORLAS Creator from O.R. Laser Technologie GmbH. The settings were 212.5W laser power, a track width of 57.5 mm, a beam diameter of 76.7 mm, a layer height of 30 mm and a scan speed of 1,200mm s 1. Both samples were printed using the aforementioned coreshell strategy with the part shapes scanned or extruded with two perimeter tracks and the core filled by a unidirectional hatching strategy. As can be seen in Figure 1(b) resp. Figure 2(b) the slicers used for each part utilized a slightly different strategy to generate the hatching. The LPBF part was sliced using the ORLas Suite Version 5.5.0.0 (O.R. Lasertechnologie GmbH, 2018) and generated distinct single scan paths for the exposure of the core of the part, whereas the FFF part was sliced using Slic3r 1.3.0 (Ranellucci and Lenox, 2018), which generated a single, connected path to fill up the core of the part. The LPBF print path contains considerably more single paths for this reason but also a longer total path length per area because of the significantly lower hatching distance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003246_robosoft48309.2020.9115976-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003246_robosoft48309.2020.9115976-Figure4-1.png", + "caption": "Fig. 4: Schematic representation of the experimental setup. Obstacle 1 and 2 are a cylindrical tube (diameter: 40 mm, height: 410 mm), and a plexiglass sheet (300 x 500 x4 mm3). In the setup, it is r = 40 mm, and w = 80 mm from O that denotes the center of the module.", + "texts": [ + " For each event detected, we store the information about the expected height, the measured one, and the changes in the x\u2013 and y-values of the magnetometers. In this section we describe the results for the following scenarios: i) the system is exploring a space where no obstacles or supports are placed in the surroundings, ii) when a vertical obstacle smaller in size than the system is placed in correspondence of the magnetometer, and ii) when the system is moving in the proximity of a vertical wall placed at 180\u25e6 with respect to the magnetometer. Fig. 4 shows a top view of the experimental setup. Here, we report the results of the proposed approach for a single module. We assembled a module to the base and connected the cables to the motors. Starting from the initial relaxed configuration, we have pulled all the cables at constant speed in order to achieve a partially compressed state in which the top of the module is parallel to the ground, and then record the position of the magnet. This pose has been set as home pose. The tests described in the followings have started from this configuration", + " 5 shows the reading of the sensors together with the fitting of the model on the acquired data. The values are calculated as the mean over a sequence of 10 nutation performed with constant bending angle. As expected, by increasing the shortening of the cables, the module can reach farther location and thus explore a larger area. C. Identification of a Vertical Support Following the results of the previous set of experiments, we have tested the system for the detection of a vertical support placed as depicted in Fig. 4 at distance R from the rotation axis of the module, while the system is performing nutation at constant speed. Since no support was detected, we gradually performed further bending until reaching the total shortening of the cables of 16 mm, thus stopping at 819 Authorized licensed use limited to: University of Exeter. Downloaded on June 24,2020 at 04:27:35 UTC from IEEE Xplore. Restrictions apply. the maximum achievable angle. This was required since a single module can explore only a limited portion of the space", + " By repeating the measurements over a series of 5 nutation, and by counting the occurrence of mismatch, it is possible to identify the location of the obstacle within \u00b160\u25e6 with respect to the position of the magnet that corresponds to the origin. D. Identification of a Impassable Wall A last series of experiment was performed to test the possibility to distinguish between small and large objects. In the specific, we tested the system in identifying the location of an impassable wall at distance, d, and places as shown in Fig. 4. As in the previous case, we first perform a series of nutation with a gradually change in the bending angle until a first mismatch with the model is detected. In this case, since the obstacle was placed ad smaller distance from the rotation axis of the module, the total shortening of the cables was less (11 mm). Fig. 7 shows the results one of the nutation (solid line) compared to the average value of the ones performed with the same configuration but in absence of any obstacle (dashed line). As it is possible to notice in the figure, there is a clear change in the readings in the area between 100\u25e6 and 280\u25e6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002115_090504-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002115_090504-Figure5-1.png", + "caption": "Fig. 5. (color online) Lu\u0308 chaotic attractor.", + "texts": [ + "055]. After 500 time steps, the system state trajectories are given in Fig. 2, and the tracking error trajectories are shown in Fig. 3. The control error trajectories are given in Fig. 4. It can be seen that the proposed method can make the chaotic system track the desired objective. We consider the Lu\u0308 system,[25,31,32] which is described by x\u0307 = f +gu, (43) where f = a(x2\u2212 x1) \u2212x1x3 + cx2 x1x2\u2212bx3 , and g = diag(5,5,5). When a = 36, b = 3, and c = 20, the internal system Eq. (43) is shown in Fig. 5. Let the desired trajectories be \u03b8 = [2;\u22122;0.5]. The initial weights of critic and action networks are selected in (\u22120.5,0.5). Based on the proposed method, the weights of critic and action networks converge to WJ = [0.0093;\u22120.0454;0.0489] and Wv = [\u22120.0196, \u22120.0165, 090504-5 \u22120.0324;\u22120.0848, \u22120.0899, 0.0799;\u22120.0520, 0.0804, \u22120.0260;\u22120.0753, 0.0889, \u22120.0777;\u22120.0632, \u22120.0018, 0.0560;\u22120.0520, \u22120.0021, \u22120.0220]. After 400 time steps, the chaotic system trajectories are shown in Fig. 6. The tracking error trajectories and control error trajectories are given in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002656_icamechs.2016.7813415-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002656_icamechs.2016.7813415-Figure5-1.png", + "caption": "Fig. 5. Overview of mowing mechanism", + "texts": [ + "0 [kgf], including the battery Its width should be less than 350 [mm] so that it can mow in narrow spaces We have designed a mowing system that satisfies these specifications, as described in the next section. III. DEVELOPMENT OF TRIMMER-TYPE MOWING SYSTEM First, an overview of the proposed mowing system is shown in Fig. 4. The system consists of a mowing unit, a driving unit, and a control unit. The specifications of the mowing system are listed below. An overview of the mowing unit is shown in Fig. 5. The top part of the figure shows DC motor that rotate, whereby power is transmitted to a cam by spur gear. The cam pushes a reciprocating bar, and the cutting edge moves sideways. As a result, the moving edge and fixed edge shear the grass. The driving mechanism remodels research and development crawler (FRIGO [8]). The drive performance of a crawler on an irregular ground is high. Crawler belts transmit the output of DC motors to the left and right sprockets via timing belts (Fig. 6). We use a high-output motor, set perpendicularly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001333_ilt-10-2011-0080-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001333_ilt-10-2011-0080-Figure1-1.png", + "caption": "Figure 1 TSUJ assembly", + "texts": [ + " (2000a, b, 2005), theoretically and experimentally. The kinematic and dynamic properties thereof were also theoretically investigated by Mariot and K\u2019nevez (1999, 2002), Mariot et al. (2004) and K\u2019nevez et al. (2001). Serveto et al. (2008) carried out the theoretical and experimental studies on the axial force of the tripod joint. However, the tripod sliding universal joint (TSUJ) has been presented in recent years in order to transmit torque larger than before (Wang et al., 2009; Wang and Chang, 2009a). It can be seen from Figure 1(b) that in general TSUJ consists of the input shaft, three slide rods, three joint bearings, and the tripod with the output shaft. Compared with the tripod universal joint of roller type, it is through three slide rods instead of three spherical or cylindrical rollers that its input shaft and tripod link up. For lubricant does not enter smoothly the gap between the slide rods and the holes of the input shaft, the severe wear (Figure 2) presents itself which restrains its popularization and utilization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003526_j.asoc.2020.106729-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003526_j.asoc.2020.106729-Figure1-1.png", + "caption": "Fig. 1. 3D assembly structure and hand frame {Ot } of the flexible hand.", + "texts": [ + " The main limitation of these pproaches is that the prediction accuracy of the pose of the rasped object is poor, which provides only a certain number of rasp directions to the grasp configuration. Considering the probem of grasping unknown objects, specifically novel objects that re observed for the first time through 2D images, an effective nd accurate estimation of grasp configuration is crucial for grasp ynthesis. . Flexible hand with a variable palm and soft fingers In this work, we develop a novel underactuated flexible hand, hich includes a variable palm and four soft fingers, as shown n Fig. 1. The variable palm provides a flexible grasp range and supporting skeleton for the flexible hand, which is composed f a metacarpal mechanism and a palmar face with a soft plane Fig. 1). .1. Variable palm Fig. 2 presents the kinematic model of the metacarpal mechnism of the variable palm, which describes the palm-changing echanism by using four links (i.e. lk1, lk2, lk3 and lk4) for the lexible hand. A translational slider (i.e. the translation distances re ds1, ds2, ds3 and ds4) exists on each link that can be used as a prismatic joint. Links lk2 and lk4 can rotate (i.e. the rotation angles re \u03b8p2 and \u03b8p4, respectively) around the coordinate centre Ot . he initial configuration of the metacarpal mechanism is that the our links are perpendicular to one another (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003911_s42417-020-00269-4-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003911_s42417-020-00269-4-Figure5-1.png", + "caption": "Fig. 5 A pair of helical compression springs for one slot", + "texts": [ + " The values for the model parameters are shown in Table\u00a01 [20]. (1) \u23a1\u23a2\u23a2\u23a2\u23a3 I1 0 0 0 0 I2 0 0 0 0 I3 0 0 0 0 0 \u23a4\u23a5\u23a5\u23a5\u23a6 \u23a7\u23aa\u23a8\u23aa\u23a9 ?\u0308?1 ?\u0308?2 ?\u0308?3 Z\u0308 \u23ab\u23aa\u23ac\u23aa\u23ad + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 C12 \u2212C12 0 0 \u2212C12 C12 + C23 \u2212C23 0 0 \u2212C23 C23 0 1 \u2212 z \ud835\udeff sign \ufffd ?\u0307?2 \u2212 ?\u0307?1 \ufffd \u2212 \ufffd 1 \u2212 z \ud835\udeff sign \ufffd ?\u0307?2 \u2212 ?\u0307?1 \ufffd\ufffd 0 1 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a7\u23aa\u23a8\u23aa\u23a9 ?\u0307?1 ?\u0307?2 ?\u0307?3 Z\u0307 \u23ab\u23aa\u23ac\u23aa\u23ad + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 k12 \u2212k12 0 \u2212 Fmax \ud835\udeff \u2212k12 k12 + k23 \u2212k23 Fmax \ud835\udeff 0 \u2212k23 k23 0 0 0 0 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a7 \u23aa\u23a8\u23aa\u23a9 \ud835\udf031 \ud835\udf032 \ud835\udf033 Z \u23ab \u23aa\u23ac\u23aa\u23ad = \u23a7\u23aa\u23a8\u23aa\u23a9 \ud835\udf0f 0 0 0 \u23ab \u23aa\u23ac\u23aa\u23ad 1 3 In general, a clutch disc has a set of helical compression springs. A single pair of springs is shown in Fig.\u00a05, with cylindrical springs of two different diameters. Four pairs of springs are shown in Fig.\u00a06 inside the slots of the damper, as commonly found in clutch discs. The equivalent stiffness for these springs mounted in parallel results in stiffness k12. Tuning tests are carried out in the automotive industry to evaluate the vehicle by measuring the angular acceleration of the drivetrain components under five conditions (idle, creeping, drive, coast and engine start/stop). In this study, the drive condition was analyzed, which is the most common and severe condition of the powertrain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003445_j.procir.2020.05.186-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003445_j.procir.2020.05.186-Figure3-1.png", + "caption": "Fig. 3. Tool path of the touch probe in hyperCAD\u00ae-S (black lines: Rapid traverse (G0), grey lines: working process (G1)", + "texts": [ + " The numbers of the applied points are listed in table 1. The calculation of the measuring point positions according to the Hammersley sampling strategy was done in MATLAB\u00ae (The MathWorks Corporation, Natick, USA). In Figure 2 the realization of the Hammersley sampling strategy of the contour for the free form surface is shown. After the definition of the measuring points depending on the considered geometric feature, the measuring cycle was implemented in the CAM-program hyperCAD\u00ae-S (OPEN MIND Technologies AG, Wessling, Germany, figure 3). Therefore, the \u201d3D-point-probing\u201d cycle and a touch trigger probe (Infrared Probe 25.41-HDR, m&h Inprocess Messtechnik GmbH, Walburg, Germany) were used. In this cycle, single measuring points can be chosen and a tolerance can be defined. Parts, that will never meet the requirements, because of no possibility of fitting them, may be scrap. The tolerance-definition is an abort criterion for the post-processing of a part. In table 2 the results of an exemplary measurement log for the circular ring (floor surface) in z-direction are shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000614_icage.2014.7050162-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000614_icage.2014.7050162-Figure2-1.png", + "caption": "Figure 2 Winding arrangement of 2 pole 3 phase induction motor", + "texts": [ + " H\u03a8=1 for E\u03a8 > +HB\u03a8 (1) H\u03a8= -1 for E\u03a8 > -HB\u03a8 (2) where 2HB\u03a8 is the total hysteresis bandwidth of the controller. The actual stator flux is constrained within the hysteresis band and tracks the command flux. The torque controller has two levels of digital output according to following equation. HT=1 for ET > +HBT (3) HT= -1 for ET > -HBT (4) HT= 0 for -HBT < ET < + HBT (5) The lookup table for conventional DTC scheme is given below. TABLE 1 LOOKUP TABLE FOR CONVENTIONAL DTC The winding arrangement of a two-pole, three-phase \u2013 Y connected induction machine is shown in Figure 2. The stator windings of which are identical, sinusoidally distributed in space with a phase displacement of 120 0 , with equivalent turns and resistance rs In order to apply generalised machine theory to polyphase induction machines, it is essential to have d,q axis fixed on the stator. Since the 3-phase winding A,B,C on the stator and d,q axes are stationary with respect to each other the transformation matrix from ABC to dq or vice versa should contain constant coefficients. On expanding the above matrix transformation equation , ids= Im cos(\u03c9t+\u03b1) and iqs=Im sin(\u03c9t+\u03b1) This indicates that d,q axes currents are also functions of time and are displaced each other by a phase angle 90 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001859_2016-01-0814-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001859_2016-01-0814-Figure2-1.png", + "caption": "Figure 2. In-cylinder pressure acting on engine head. Part of picture from [19]", + "texts": [ + " The proposed solution falls among those based on the measurement of the force produced by combustion gasses and was conceived to allow an easy installation, be directly correlated with the pressure inside the cylinder, have no limits related to engine operating conditions and have a simple post processing. Conceptually, the pressure acts on the combustion chamber surface and generates a force that is elastically transferred through the cylinder head to the head bolts or studs. Therefore, by measuring this force (e.g. by a strain washer), it is possible to obtain a direct reading of the variations of the combustion chamber pressure (Figure 2). Since the installation of the strain washer needs a stud longer than those originally present on the engine, all the studs have been substituted with longer ones to guarantee a uniform load distribution. The experimental activity was carried out on a two-stroke spark ignition single cylinder engine (Table 2) with a displacement of 300 cm3 and a low-pressure direct injection (LPDI) system [20]. The engine was equipped with a dynamic pressure sensor installed on the engine head in order to measure the in-cylinder pressure curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003694_ecce44975.2020.9235404-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003694_ecce44975.2020.9235404-Figure2-1.png", + "caption": "Fig. 2. Flux paths in an FSPM machine.", + "texts": [ + " The effects of skew on the thermodynamic performance and electromagnetic performance of the integrated motor-compressor are investigated in Section III and Section IV, respectively. Section V provides the conclusion. 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 5545 Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 22:42:22 UTC from IEEE Xplore. Restrictions apply. The topology of the integrated flux-switching motorcompressor is provided in this section. The motor topology of the integrated motor-compressor is the FSPM machine. Fig. 2 shows the flux paths of an FSPM machine, and the flux in the salient rotor poles are bidirectional. FSPM machine has the advantage of robust rotor structures as the magnets are sandwiched between the stator cores. As a result, there are no magnets retainment issues with respect to other PM machines with rotor magnets. The salient rotor poles of the FSPM machine are shaped into airfoils to provide compression function. NACA 65(4)-421 blade is selected for the airfoil-shaped rotor accounting for the tradeoffs of blade thickness between the motor and compressor [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.43-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.43-1.png", + "caption": "Figure 9.43 Outward elbow (E) and inward elbow (E\u2032) poses of the manipulator.", + "texts": [ + "707) can be manipulated as follows: cER = [d4eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 u\u03032e\u2212u\u03033\ud835\udf033 e\u2212u\u03032\ud835\udf032][d6eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 eu\u03033\ud835\udf035 u2] \u21d2 cER = d4d6eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 u\u03032eu\u03033\ud835\udf035 u2 = d4d6eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 u\u03032[u2c\ud835\udf035 \u2212 u1s\ud835\udf035] \u21d2 cER = d4d6eu\u03032\ud835\udf032 eu\u03033\ud835\udf033 u3s\ud835\udf035 = d4d6eu\u03032\ud835\udf032 u3s\ud835\udf035 cER = d4d6(u3c\ud835\udf032 + u1s\ud835\udf032)s\ud835\udf035 (9.711) Upon substituting Eqs. (9.661) and (9.662), Eq. (9.711) becomes cER = \ud835\udf0e2d4d6(u1x3 \u2212 u3x1)s\ud835\udf035\u2215 \u221a x2 1 + x2 3 (9.712) At this point, it is also to be noted that u\u03032x = u\u03032(u1x1 + u2x2 + u3x3) = u1x3 \u2212 u3x1 (9.713) Hence, cER is finally expressed as cER = x\u0303Ex = (\ud835\udf0e2u\u03032x)d4d6 \u221a 1 \u2212 \ud835\udf092 5\u2215 \u221a x2 1 + x2 3 (9.714) As for the role of cER in determining the pose of the arm, it is illustrated in Figure 9.43. This figure implies that the arm assumes the outward elbow pose if the senses of the cross products cER = x\u0303Ex and c2x = u\u03032x are equal. According to Eq. (9.714), this equality can be realized if \ud835\udf0e2 = + 1. Otherwise, i.e. if \ud835\udf0e2 = \u2212 1, the arm assumes the inward elbow pose. u3 (0)\u2192 E\u02b9 E\u02b9 E O S O S R R E u1 (0)\u2192 u2 (0)\u2192 u2 (0)\u2192 u1 (0)\u2192 u3 (0)\u2192 326 Kinematics of General Spatial Mechanical Systems (c) Third Kind of Multiplicity The third kind of multiplicity is associated with the sign variable \ud835\udf0e7 that arises in the process of finding \ud835\udf037" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002262_978-3-319-27333-4-Figure3.7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002262_978-3-319-27333-4-Figure3.7-1.png", + "caption": "Fig. 3.7 Lubrication unit", + "texts": [ + " The measured data of oil temperature and speed obtained during the experimentation in analog form is converted into digital form and preprocessed on the instrumentation card. Then this data is serially transmitted to PC through signal output cable 1 and 2. The data is received, displayed, and stored on the PC using Lab view-based software WINDUCOM 2005. Lubrication unit is made of a metallic tank with a motor and pump, by pass valve, control valve, pressure gauges, flow meter, inlet, and delivery pipe (Fig. 3.7). The discharge of the oil is controlled by means of gate valves provided. Two oil inlets on journal bearing are provided with a pipe, its flow is regulated by two gate valves, and pressure at that point is indicated by two pressure gauges fixed in line. An oil sump is provided beneath the bearing for collecting the used oil, which flows into metallic tank for its re-circulation. Figure 3.8 shows the main structure of the journal- bearing test rig. 3.3 Experimental Procedure 30 3 Performance Parameters 31 The test rig developed has been further modified to take higher loads and new test rig has been developed under AICTE Research Project (letter No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001825_2016-01-1092-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001825_2016-01-1092-Figure2-1.png", + "caption": "Figure 2. Inertia test stand with gearbox enclosed in isolation box;", + "texts": [ + " The easily interchangeable mounting plate is designed to allow gearshift tests or drag torque measurements of different transmissions. The clutch plate is inserted on the gearbox input shaft, in order to keep the realistic moment of inertia of the gearbox primary part. and flexible coupling, 6 - Strain gauge with telemetric device, 7 - Drive shaft, 8 - Tested gearbox, 9 - Safety brake, - Sensor for speed of rotation, - Torque sensor, - Temperature sensing. To obtain wide ranges of gearbox temperatures, the gearbox was enclosed into an isolation box (see Figure 2). To simplify the form of isolation box and to be sure that there is no tension in shift cables (whatever the gearbox temperature would be) the shift robot was disconnected and the speeds were shifted manually. The following parameters were measured to obtain the no-load losses in the gearbox: \u2022 Temperature of the gearbox. \u2022 Temperature of the bearing housing. \u2022 Torque. \u2022 Speed of rotation of the flywheel. \u2022 Speed of the rotation of gearbox input shaft. Additionally sensed values are not listed here, because they serve only for monitoring right functioning of the experimental stand" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003929_s40430-020-02785-6-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003929_s40430-020-02785-6-Figure3-1.png", + "caption": "Fig. 3 Structure diagram of the WDSB", + "texts": [ + " When the structure and working conditions are known, the mass, stiffness, and damping coefficient in the above equation are determined. Supposing m1 = 1100\u00a0kg, m2 = 30\u00a0kg, \u03c9 = 20\u00a0Hz, F0 = 3kN, k1 = 1\u00d7107\u00a0N/m, k2 = 1\u00d7109\u00a0N/m, c1 = 0.4 \u00d7 105\u00a0N/(m/s), bring into equations to calculate the journal response with c2. WDSB\u2019s journal response increases along with the increase in c2 following a decrease. When c2 is 8.1 \u00d7 106\u00a0N\u00b7s/m, WDSB\u2019s journal response is the smallest, about 27.7% lower than that of CWSB. In view of the independent design of load carrying and vibration reduction, a WDSB with ISFD (as shown in Fig.\u00a03) is designed. \u201cS\u201d shape gaps are set in the CWSB\u2019s bushing. According to the carrying characteristics of WSB, the bearing\u2019s attitude angle is very small, about 5\u00b0 ~ 20\u00b0, and the carrying area is concentrated in the range of 20\u00b0 ~ 30\u00b0 at the bottom of the bearing [13]. Therefore, the scheme only has two \u201cS\u201d-shaped elastic structures at the bottom of the bearing, which are arranged along the circumference and symmetrical along the vertical centerline. In order to analyze the effects of the structure on the damping effect, three structural parameters are defined namely oil film clearance \u03b4, inner flange height s, distribution angle \u03b8 of \u201cS\u201d shape springs, as shown in Fig.\u00a03. The \u201cS\u201d shape springs divide the bushing into two parts, inner flange and outer flange. Oil is enclosed in the \u201cS\u201d shape gaps by end caps, and an oil chamber is arranged on the end cover to store oil. When the bearing is operating, the \u201cS\u201d shape springs will come a rhythmic stretching deformation under the action of the excitation force. The oil is sucked and discharged along the axial direction in the \u201cS\u201d shape gaps, leading to a squeeze effect. At the same time, the \u201cS\u201d spring side wall and the oil cover are relatively running, and a piston/dash-pot effect is formed in the gap between them" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001080_s00170-013-4779-2-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001080_s00170-013-4779-2-Figure1-1.png", + "caption": "Fig. 1 Single-joint robotic system concept", + "texts": [ + " The convergence analysis is established for both adjustable parameters inside FREN and the tracking error. The experimental system with a 7-DOF robotic arm (Mitsubishi PA10) is carried out to verify the performance of the proposed algorithm. Most robotic systems especially in the industrial sector are velocity-mode controllers. To simplify, in this work, we consider that the control effort which will be fed to the robot is the velocity command denoted by u and the output is the joint position denoted by x. This concept can be illustrated in Fig. 1 as a single-joint robotic system. For the system based on a digital computer controller, generally, we presume that the relation between joint position and velocity command can be defined by the formulation in a class of nonaffine discrete-time systems as x(k + 1) = fN(\u03bc(k), u(k)), (1) where fN(\u00b7) is an unknown nonlinear function, \u03bc(k) = [x(k) x(k\u22121) \u00b7 \u00b7 \u00b7 x(k\u2212n) u(k\u22121) u(k\u22122) \u00b7 \u00b7 \u00b7 u(k\u2212m)]T , and k stands for the time index. Parameters n and m denote as the order of system\u2019s output and input, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002569_iceee.2016.7751210-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002569_iceee.2016.7751210-Figure2-1.png", + "caption": "Fig. 2. Definition of forces, moments and velocity components in a body fixed frame [7].", + "texts": [ + " The experimental results are presented using a Radian Pro Aircraft [1]. Lateral dynamics includes roll, yaw and sideslip motions of aircraft. The yaw motion is achieved by deflecting a flap on the rudder, and sideslip motion is entered by lowering a wing and applying exactly enough opposite rudder so the aircraft does not turn. Yb and Zb represent aerodynamics force components, \u03d5 and \u03b4a represent the orientation of the aircraft in the earth-axis 978-1-5090-3511-3/16/$31.00 2016 IEEE system and aileron deflection angle respectively. Fig. 2 shows the forces, moments and velocity components in the body fixed frame of an aircraft. Roll control system is shown in Fig. 1, where L,M and N represent the aerodynamic moment components, the term p, q and r represent the angular rates components of roll, pitch and yaw axis and the term u, v and w represent the velocity components of roll, pitch, and yaw axis. In the model devlopement, we asume that the aircraft is in steady cruise with constant altitude as well as constant velocity, that a change in the pitch angle does not change de the speed of the aircraft and that the reference flight conditions are symmetric with propulsive forces constant [8] [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure7.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure7.1-1.png", + "caption": "Figure 7.1 A typical serial manipulator with six revolute joints.", + "texts": [ + " In a special purpose manipulator, the end-effector may be one of the devices such as a pen, a screwdriver, a drilling tool, a welding tool, a painting gun, a multi-finger gripper, etc. Considering a link or a joint with its position in the kinematic chain, it is called proximal if it is closer to the base and distal if it is further away from the base. Kinematics of General Spatial Mechanical Systems, First Edition. M. Kemal Ozgoren. \u00a9 2020 John Wiley & Sons Ltd. Published 2020 by John Wiley & Sons Ltd. Companion Website: www.wiley.com/go/ozgoren/spatialmechanicalsystems A typical serial manipulator with six revolute joints is illustrated in Figure 7.1 showing its side and top views. The axes of the joints are represented by the unit vectors n\u20d71, n\u20d72, \u2026, n\u20d76. The kinematic features of a serial manipulator can be described conveniently according to the Denavit\u2013Hartenberg (D\u2013H) convention. The D\u2013H convention is based on the joint axes and the common normals between them. However, there have so far been different formulations proposed for the D\u2013H convention. Aside from the notational differences, these formulations differ from each other basically by the indexing rules" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure5-1.png", + "caption": "Figure 5. Overall view and a cross-sectional diagram of the gear with different passive roller shapes", + "texts": [ + " In this gear structure, the rotational axis of each passive roller is aligned continuously on the circular trajectory around the center shaft of the whole gear mechanism. Each rotational axis of a passive roller is independent. This gear mechanism eliminates the need for multiple layers of passive rollers like the gear with conical passive rollers shown in the previous chapter for sliding with an omnidirectional gear with an intermittent teeth structure. The third gear mechanism we examined has a structure resembling an ordinary commercial omnidirectional wheel with the barrel shape of passive rollers. A schematic diagram of this gear is shown in Fig. 5. In this gear structure, the passive rollers in one group between the supporting plates have a common rotational axis, while the rotational axes of passive rollers are aligned intermittently in a circular direction around the center shaft of the whole gear structure. Accordingly, this gear structure has multiple layers of passive rollers with some offset in a circular direction for continuous contact with the driven omnidirectional gear. Compared with the gear structure with flat passive rollers shown in Fig. 4, the gear structure shown in Fig. 5 can be more easily miniaturized because there are fewer rotational shafts for passive rollers. As shown above, each gear structure with passive rollers has drawbacks and advantages. Accordingly, we should choose an optimum structure according to the application of a system comprising a gear with passive rollers and an omnidirectional driving gear mechanism. The proposed gear mechanism with passive rollers has a completely different structure from previous research, whereby the direction of passive rotational axes of its small rollers was aligned with the active rotational axis of the whole gear structure [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003595_icuas48674.2020.9214014-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003595_icuas48674.2020.9214014-Figure1-1.png", + "caption": "Figure 1. The ducted fan studied in this paper and its layout", + "texts": [ + " INTRODUCTION In recent years, performance requirement for unmanned aerial vehicles (UAVs) has been more and more focused on \u2018complex environment working ability\u2019. To be capable of both steady hover and high-speed flight, one preference is the ducted fan UAV. During the past decades, many classical models of ducted fan have been developed, such as iSTAR [1], HoverEye [2] and GTSpy [3]. In this paper, we study a small ducted fan UAV with typical configuration that the fan/rotor/propeller is shrouded by the duct with relatively tiny gap between the fan tip and the duct wall (See Fig.1). Taking use of the airflow exhausted from the fan, fixed aerodynamic flaps are placed to generate a moment that counteract the fan torque, while deflectable control surfaces are set to control the aircraft attitude. This particular compact layout allows the aircraft working safely in narrow spaces like warehouse. At hover, a ducted fan performs like a rotorcraft. However, when levelly flying at high-speed, a ducted fan is subject to a considerable lift and features more like a fixed-wing airplane, which leads to less power consumption than hover", + " It is worth notice that the NN evolution (18) only leads to the convergence of the state predictor 4x\u0302 , which is proved by Lemma 1. Generally, independent convergence of weights 1W\u0302 and 2W\u0302 cannot be guaranteed. This is obviously shown in Fig.4(g) and Fig.4(h) in the time of 10s to 15s. During this period, overshoot happens on the matching of 22f while certain lag occurs on the that of 22g in order to maintain a good the tracking performance of the relevant state. In this section, experimental results are presented from practical flight tests. The ducted fan prototype shown in Fig.1 is mounted with IMU and GPS to measure and estimate the system states. Data fusion, state estimation, communication and control scheme modules are coded into the airborne processor. The NN weights are chosen the same as (43). The transition trajectory is similar to that in Section IV described by (48) (i.e. fixed altitude and constant horizontal acceleration). Unlike the training approach in the simulation, as for the 1336 Authorized licensed use limited to: Middlesex University. Downloaded on October 18,2020 at 14:25:56 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002135_amc.2016.7496392-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002135_amc.2016.7496392-Figure4-1.png", + "caption": "Fig. 4. Definition of coordinate frames", + "texts": [ + " |ri| \u2264 ri,max 2) Structural solvability (inverse kinematics problem): Hands and feet can be placed on the holds 3) Force direction: Force vectors are inside of the hold cones. 4) Holding power: Holding forces are smaller than the maximum forces. |Fi| \u2264 Fi,max 5) Torque constraint: Maximum joint torques are smaller than maximum torques of the robot, \u03c4i,j \u2264 \u03c4ij,max A. Idea of reachability and regions Consider the reachability of climbing robot. Where the word reachability in this paper does not mean the reachability in control theory. It means whether hands and feet can reach holds or not. First, Fig.4 expresses the coordinate frames defined on climbing robotin 3D space. It is assumed that the gravity force mg is given at the center of mass(COM) of the robot. Moreover, we define five coordinate frames on COM and four limbs in this paper. Second, Fig.5 shows the movilable region corresponding to the prototype robot specification and 2D space. It is assumed that the robot cannot cross the hands or feet each others and the righ thand, left hand, riight foot, and left foot have their own mobilable regions, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003542_icist49303.2020.9202213-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003542_icist49303.2020.9202213-Figure3-1.png", + "caption": "Fig. 3. Aerosonde UAV and its Simulink block", + "texts": [], + "surrounding_texts": [ + "The Aerosonde block[20], a near real UAV model which is modified from the AeroSim Blockset 6-DoF aircraft model, is selected as the simulated UAV airframe. Aerosonde UAV, with gross aircraft mass of 13.5kg, a wing span of 2.8956m, a wing area of 0.55m2, an airspeed range of 15~50m/s, a sideslip angle range of -0.5~0.5rad, an angle of attack range of -0.1~0.3rad, has good aerodynamic performance and maneuverability(refer to [20] for detailed aircraft configuration) The control system of Aerosonde UAV is a collection of PID controller which are used to control surface deflections and throttle to meet desired conditions without human intervention, including roll attitude autopilot, airspeed hold controller and altitude hold control system. The gain values for the controllers were obtained through experience and are presented in Table I. TABLE I. GAIN VALUES FOR THE CONTROLLERS[20] Control Surface Control Type Gain Value Bank-to-Ailerons Proportional 0.5 180 Integral 0.05 180 Control Surface Control Type Gain Value Bank-to-Rudder Proportional 0.3 180 Integral 0.03 180 Airspeed-to-elevator Proportional 0.08 Integral 0.01 Derivative 0.3 Altitude-to-throttle Proportional 0.15 Integral 0.000001 Derivative 0.01" + ] + }, + { + "image_filename": "designv11_34_0000855_kem.579-580.300-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000855_kem.579-580.300-Figure2-1.png", + "caption": "Fig. 2 The solving process of cylindrical gear turned angle", + "texts": [ + " ))1/(1),1/(1( 2 2 22 1 2 \u2212\u2212\u2264 mnmnmink . (3) Modeling of Cylindrical Gear. The tooth profile of non-circular gear are enveloped by the tooth profile of cylindrical gear, so the tooth profile of cylindrical gear should has higher accuracy. This paper fit the points on the tooth profile with cubic spline curve, and scan a straight or helical tooth based on the generated involute profile, then array all teeth, the generated involute profile and cylindrical gear are shown in Fig. 1. Realization of Pure Rolling. As shown in Fig. 2, make four-order elliptic gear with concave as an example, coordinate system )-( xyoS is fixed with gear billet and )-( 111 yxoS is mobile with cylindrical gear but without rotation. When the pith circle roll from A to B , the turned \u03d5 of cylindrical gear around its center consists of the appendant angle 1 \u03d5 caused by the rotation around the center of polar coordinates and the angle 2 \u03d5 caused by the pure rolling arc length AB S . The coordinates of B point in coordinate system )-( 111 yxoS is as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.2-1.png", + "caption": "Figure 9.2 Line diagrams that show the kinematic details.", + "texts": [ + " Its generic name comes from the acronym PUMA that stands for \u201cProgrammable Universal Manipulation Arm.\u201d It was developed originally by Victor Scheinman at the Unimation Robot Company. Typical versions of Puma manipulators are the Puma-560 and Puma-600. A Puma manipulator comprises six revolute joints. So, it is designated symbolically as 6R or R6. Its wrist is spherical. Its joint axes are also shown in Figure 9.1 together with the relevant unit vectors. Its kinematic details (the joint variables and the constant geometric parameters) are shown in the line diagrams in Figure 9.2. The line diagrams comprise the side view, the top view, and two auxiliary views that show the joint variables that are not seen in the side and top views. The significant points of the manipulator are as follows: Kinematics of General Spatial Mechanical Systems, First Edition. M. Kemal Ozgoren. \u00a9 2020 John Wiley & Sons Ltd. Published 2020 by John Wiley & Sons Ltd. Companion Website: www.wiley.com/go/ozgoren/spatialmechanicalsystems O: Center point or neck point (origin of the base frame) S: Shoulder point, E: Elbow point, R: Wrist point, P: Tip point (a) Joint Variables \ud835\udf031, \ud835\udf032, \ud835\udf033, \ud835\udf034, \ud835\udf035, \ud835\udf036 They are the rotation angles about the joint axes. They are shown in Figure 9.2. (b) Twist Angles \ud835\udefd1 = 0, \ud835\udefd2 = \u2212\ud835\udf0b\u22152, \ud835\udefd3 = 0, \ud835\udefd4 = \ud835\udf0b\u22152, \ud835\udefd5 = \u2212\ud835\udf0b\u22152, \ud835\udefd6 = \ud835\udf0b\u22152 (c) Offsets d1 = 0, d2 = OS, d3 = 0, d4 = ER, d5 = 0, d6 = RP Kinematic Analyses of Typical Serial Manipulators 235 The nonzero offsets are given the following special names. d2: Shoulder offset, d6: Tip point offset d4: Front arm length (i.e. the overall length of Links 3 and 4) (d) Effective Link Lengths b0 = 0, b1 = 0, b2 = SE, b3 = 0, b4 = 0, b5 = 0 The only nonzero effective link length is b2. It is called the \u201cupper arm length\u201d (i", + "9) C\u03024 = C\u0302(0,4) = C\u0302(0,3)C\u0302(3,4) = eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 e\u2212u\u03031\ud835\udf0b\u22152eu\u03031\ud835\udf0b\u22152eu\u03033\ud835\udf034 \u21d2 C\u03024 = eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 eu\u03033\ud835\udf034 (9.10) C\u03025 = C\u0302(0,5) = C\u0302(0,4)C\u0302(4,5) = eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 eu\u03033\ud835\udf034 e\u2212u\u03031\ud835\udf0b\u22152eu\u03033\ud835\udf035 \u21d2 C\u03025 = eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 e\u2212u\u03031\ud835\udf0b\u22152 (9.11) C\u03026 = C\u0302(0,6) = C\u0302(0,5)C\u0302(5,6) = eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 e\u2212u\u03031\ud835\udf0b\u22152eu\u03031\ud835\udf0b\u22152eu\u03033\ud835\udf036 \u21d2 C\u03026 = C\u0302 = eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 eu\u03033\ud835\udf036 (9.12) In Eq. (9.9), \ud835\udf0323 is introduced as a combined joint variable. It is defined as follows: \ud835\udf0323 = \ud835\udf032 + \ud835\udf033 (9.13) The angle \ud835\udf0323 is also shown in Figure 9.2. It has a particular significance because it indicates the declination of the front arm with respect to the third axis of the base frame. 236 Kinematics of General Spatial Mechanical Systems (c) Location of the Wrist Point with Respect to the Base Frame r\u20d7 = r\u20d7OR = r\u20d7OS + r\u20d7SE + r\u20d7ER = d2u\u20d7(2) 3 + b2u\u20d7(2) 1 + d4u\u20d7(4) 3 (9.14) The vector Eq. (9.14) implies the following matrix equation in the base frame. r = r(0) = d2u(2\u22150) 3 + b2u(2\u22150) 1 + d4u(4\u22150) 3 (9.15) Equation (9.15) can be manipulated as shown below", + " r = d2C\u0302(0,2)u(2\u22152) 3 + b2C\u0302(0,2)u(2\u22152) 1 + d4C\u0302(0,4)u(4\u22154) 3 \u21d2 r = d2eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 e\u2212u\u03031\ud835\udf0b\u22152u3 + b2eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 e\u2212u\u03031\ud835\udf0b\u22152u1 + d4eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 eu\u03033\ud835\udf034 u3 \u21d2 r = d2eu\u03033\ud835\udf031 u2 + b2eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 u1 + d4eu\u03033\ud835\udf031 eu\u03032\ud835\udf0323 u3 \u21d2 r = eu\u03033\ud835\udf031(d2u2 + b2eu\u03032\ud835\udf032 u1 + d4eu\u03032\ud835\udf0323 u3) \u21d2 r = eu\u03033\ud835\udf031[d2u2 + b2(u1c\ud835\udf032 \u2212 u3s\ud835\udf032) + d4(u3c\ud835\udf0323 + u1s\ud835\udf0323)] \u21d2 r = eu\u03033\ud835\udf031[u1(d4s\ud835\udf0323 + b2c\ud835\udf032) + u2d2 + u3(d4c\ud835\udf0323 \u2212 b2s\ud835\udf032)] (9.16) Equation (9.16) can also be written as r = eu\u03033\ud835\udf031 r\u2032 (9.17) In Eq. (9.17), r\u2032 = u1(d4s\ud835\udf0323 + b2c\ud835\udf032) + u2d2 + u3(d4c\ud835\udf0323 \u2212 b2s\ud835\udf032) (9.18) Note that r\u2032 is the matrix representation of the vector r\u20d7 in the link frame 1(O), i.e. r\u2032 = r(1). This fact can be verified as follows: r = r(0) = C\u0302(0,1)r(1) = eu\u03033\ud835\udf031 r(1) \u21d2 r(1) = r\u2032 (9.19) Note further that, referring to Figure 9.2, r\u2032 = u1r\u20321 + u2r\u20322 + u3r\u20323 can also be written down directly by inspection. In other words, Figure 9.2 readily implies the following coordinates of the wrist point R in the frame 1(O). r\u20321 = d4s\ud835\udf0323 + b2c\ud835\udf032 r\u20322 = d2 r\u20323 = d4c\ud835\udf0323 \u2212 b2s\ud835\udf032 \u23ab \u23aa\u23ac\u23aa\u23ad (9.20) On the other hand, the expansion of the expression in Eq. (9.16) leads to the coordinates of the wrist point R in the base frame 0(O) as follows: r(0) = r = u1r1 + u2r2 + u3r3 (9.21) r = (eu\u03033\ud835\udf031 u1)(d4s\ud835\udf0323 + b2c\ud835\udf032) + (eu\u03033\ud835\udf031 u2)d2 + (eu\u03033\ud835\udf031 u3)(d4c\ud835\udf0323 \u2212 b2s\ud835\udf032) \u21d2 r = (u1c\ud835\udf031 + u2s\ud835\udf031)(d4s\ud835\udf0323 + b2c\ud835\udf032) + (u2c\ud835\udf031 \u2212 u1s\ud835\udf031)d2 + (u3)(d4c\ud835\udf0323 \u2212 b2s\ud835\udf032) \u21d2 r = u1[(d4s\ud835\udf0323 + b2c\ud835\udf032)c\ud835\udf031 \u2212 d2s\ud835\udf031] + u2[(d4s\ud835\udf0323 + b2c\ud835\udf032)s\ud835\udf031 + d2c\ud835\udf031] + u3(d4c\ud835\udf0323 \u2212 b2s\ud835\udf032) (9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003088_tia.2020.2992579-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003088_tia.2020.2992579-Figure17-1.png", + "caption": "Fig. 17. Structure of IM and traditional LSPM machine. (a) IM. (b) Traditional LSPM machine.", + "texts": [ + " Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Both of PCLSPMV machine and DSLSPMV machine acquire self-starting capability for PMV machine by stator winding pole-changing and cascading method. In order to compare the torque density between LSPMV machine and traditional self-starting machine, the FEA model of traditional LSPM machine and IM are established. The structure of IM and traditional LSPM machine is shown in Fig.17. The torque density comparison of four machines are shown in Table VII. The LSPMV machines also show the higher torque density compared with regular commercial LSPM machine. Compared with DSLSPMV machine, the structure of PCLSPMV machine is much simpler, so a prototype of PCLSPMV machine was manufactured and tested in order to verify the analysis above. The main parameters are shown in Table VII. Different from dimensions in section II, the PM width is 3.9mm for prototype. The picture of prototype and test bench are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001019_ma5000738-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001019_ma5000738-Figure2-1.png", + "caption": "Figure 2. Molecular distribution factors of (a) phenyl rings ( f x, f y, fz) with E of the LPUV light direction parallel to the x-axis, and (b) ( f x\u2032, f y\u2032, fz\u2032) with tilt angle (\u03b3), when E of the LPUV light direction is inclined \u03c8 to the x-axis.", + "texts": [ + " The angular dependences (\u03b8\u2019s) of both TEY and AEY detection were obtained by rotating a sample film perpendicular (\u03d5 = 0\u00b0) and parallel (\u03d5 = 90\u00b0) to the x-axis, where the E of LPUV light was parallel to the dx.doi.org/10.1021/ma5000738 | Macromolecules 2014, 47, 2080\u221220872081 x-axis. The reorientation direction was parallel to E when the film was exposed from the z-axis (Figures 1b,2a). The three orientation factors of the phenyl rings ( f x, f y, and fz), which indicate the relative number of phenyl rings at a molecular dx.doi.org/10.1021/ma5000738 | Macromolecules 2014, 47, 2080\u221220872082 coordinate frame (x, y, z) without a tilt angle (LPUV was incident from the z-axis, Figure 2a), were evaluated using the same technique reported in the literature.33,34 Additionally, f x\u2032, f y\u2032, and fz\u2032 indicate the relative number of phenyl rings at a molecular coordinate frame (x\u2032, y\u2032, z\u2032) with molecular tilt angle (\u03b3) when the film was obliquely exposed by p-polarized UV light (Figure 2b). Using the f x, f y, and fz values, the fractional numbers of side group segments (Nx, Ny, and Nz) were estimated using the three permutations in the expression 2f i = Nj + Nk, where i, j, and k reflect the three axes. Equation 3 expresses the inplane order parameter (S3D). The details are described in the Supporting Information. = \u2212 + + S N N N N ND x y x y z 3 (3) 3.1. Molecular Orientation Characteristics of the AsCoated PMCB6M Films. Figure 3a shows the UV absorption spectra of a 170-nm thick PMCB6M film as the film is rotated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003501_5.0015762-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003501_5.0015762-Figure1-1.png", + "caption": "FIGURE 1. Fracture of Connecting Rod", + "texts": [], + "surrounding_texts": [ + "The first stage of analysis in this study was the observation of the visual observation of the connecting rod and analysis of the fracture surface. Then, the comprehensive characterization of the connecting rod material includes microstructure test, hardness test and SEM test for observation of fracture morphology. Hardness testing is carried out according to ASTM E384- 11e1 standards and Samples for metallographic analysis are prepared according to ASTM E3-11 standards." + ] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure3.8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure3.8-1.png", + "caption": "Fig. 3.8 \u03d1 rotation about an arbitrary axis \u03bb", + "texts": [ + "2 (Euler Theorem of Rotations Goldstein 1950) The general displacement of a rigid body with one point fixed is a Rotation about some axis The Euler Theorem of Rotations is equivalent as the following proposition: Lemma 3.1 In any Rotation matrix R \u2208 SO(3) there is at least one positive unit magnitude eigenvalue as \u03bb = +1. Proof Consider that a pure rotation is achieved when a rotated frame 1 has performed a rotation angle \u03d1 along a given direction. Consider also that this direction is defined by a direction vector\u03bb, as shown on Fig. 3.8. To perform a backward rotation, then the moving frame 1 must perform the negative angular rotation\u2212\u03d1 in the very same direction \u03bb. In consequence, regardless of the magnitude of the rotation angle \u03d1, the rotation axis \u03bb should have the same coordinates in either reference frame: \u03bb(0) = \u03bb(1). Hence, the rotation equation \u03bb(0) = R\u03bb(1) yields \u03bb(0) = RT\u03bb(0) or \u03bb(1) = R\u03bb(1) (3.26) Which can be written in the form of the definition of the eigenvalues \u03bbi : R\u03bb = \u03bbi\u03bb (3.27) Comparing (3.26) and (3.27) it follows that at least one of the three eigenvalues of the rotation matrix is the unit \u03bb1 = 1, and then, the motion of the frame is given solely by a rotation above \u03bb" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003248_s00170-020-05597-z-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003248_s00170-020-05597-z-Figure4-1.png", + "caption": "Fig. 4 The ball screw joint: a double nut; b single nut", + "texts": [ + " For a single ball, the axial displacement of the nut causes the ball to deform, and the axial load is balanced with the axial force caused by the ball deformation. Therefore, the relationships among them can be written as Pa \u00bc Psin\u03b2cos\u03bb \u03b4a \u00bc \u03b4n \u00fe \u03b4m\u00f0 \u00de=sin\u03b2cos\u03bb \u00f03\u00de where Pa is the axial load on the ball, and \u03b4a is the axial displacement of the nut. \u03b6 is a ratio related to the elastic deformation at m point, \u03b4m, and the elastic deformation at n point, \u03b4n, and it can be calculated by \u03b6 \u00bc \u03b4m \u03b4n \u00bc Chm Chn \u00bc manK em\u00f0 \u00de \u2211\u03c1m\u00f0 \u00de1=3 mamK en\u00f0 \u00de \u2211\u03c1n\u00f0 \u00de1=3 \u00f04\u00de For the double-nut ball screw joint as shown in Fig. 4a, by using Eq. (2), the equilibrium and compatibility conditions can be written as Fa \u00bc N ssin\u03b2cos\u03bb PA\u2212PB\u00f0 \u00de P2=3 A \u2212P2=3 0 \u00bc P2=3 0 \u2212P2=3 B \u00f05\u00de where P0 is the preload on the normal direction of a single ball, Ns is the number of the balls in a single nut, Ns = i\u03c0ds/db cos \u03bb, and i is the number of load rings of the nut. Substituting Eqs. (2)\u2013(4) into Eq. (5), Fa can be written as Fa \u00bc N ssin\u03b2cos\u03bbC \u22123=2 hn \u03b40 \u00fe sin\u03b2cos\u03bb 1\u00fe \u03b6 \u03b4a 3=2 \u2212 \u03b40\u2212 sin\u03b2cos\u03bb 1\u00fe \u03b6 \u03b4a 3=2 \" # ; \u03b40\u2265 sin\u03b2cos\u03bb 1\u00fe \u03b6 \u03b4a \u00f06\u00de where \u03b40 is the deformation at n point under the preload P0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001289_03043797.2014.1001814-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001289_03043797.2014.1001814-Figure3-1.png", + "caption": "Figure 3. E-manufacturing projects designed and fabricated by (a) and (b) collaborative team and (c) control team.", + "texts": [ + " (c) the parts must be shipped to one location (university) for assembly, (d) the product had to not weigh more than 2 kg, and (e) The deliverables of this project included the assembled product and a report. The report must include, but not limited to, (1) introduction section (2) description of the manufacturing processes for each component (3) virtual (CAD) and physical (real) pictures of individual components and the final product (assembled), (4) test results (validation) to demonstrate that the product is functional, (5) discussion section, and (6) appendix that can include all supportive documents such as communications between members. Figure 3 shows three examples that were fabricated by collaborative and control teams. Figure 3(a) shows the mechanical iris designed and manufactured by a collaborative team. Because of the complexity of the petal design, computer-based manufacturing was required. The product consists of five types of custom components, in addition to commonly available hardware. The standard hardware purchased included threaded rods/bolts, nuts, and washers. Two different processes (water jet cutting and rapid prototyping) were used to manufacture a total of 33 custom parts. As shown in Figure 3(b), a desktop fan was designed, manufactured, and assembled by a collaborative student team. Each team member was responsible for a particular task: (1) design and manufacture (rapid prototype) of the blade, (2) design and manufacture of aluminium structure components, and (3) select and purchase of the electrical motor and assembly and testing of the fan. Figure 3(c) shows a device named \u2018Wacky Wire Game\u2019 that was manufactured via laser cutting, manual machining, rapid prototyping, and manual assembly processes. The game is played by moving a handheld metal ring to the bottom of the wire while it is spinning to win the game. A buzzer sounds via the circuit in the handle if player fails, and variable speed control allows for changing difficulty. 4.1.3. Collaborative project #3: manufacturing process development In any design firm, designing a new product requires engineers to carry out \u2018benchmarking\u2019 with best-in-class companies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000381_acc.2015.7172277-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000381_acc.2015.7172277-Figure1-1.png", + "caption": "Fig 1. Cart-pendulum system", + "texts": [ + " For a given 1 , the uncertain nonlinear system (1) under input saturation (4) with the proposed controller (8) is asymptotically stable in a DOA to be maximum if there exist matrices 11 11 0 T P P , 22 0 ijkp P , 21 nx nx ijkp P , nu nx jk K , and positive scalars a i , b ijkp , 1e ijkp , 2e ijkp such that the following LMIs optimization problem is feasible: 11, 0 min LM Is 11 -(12) s.t. LM I (29) R P (30) IV. SIMULATION RESULTS In order to illustrate the effectiveness of the proposed approach, consider now the stabilization problem of the unstable nonlinear cart-pendulum system consists of an inverted pendulum rotating in a vertical plan around an axis located on a cart moving along the x-axis as shown in Fig 1. Mass of the cart (kg) M 0.5 Mass of the pendulum (kg) m 0.2 Length to the pendulumcenter of mass (m) L 0.3 Viscous friction coefficient of the pendulum (Nsrad-1) p b 0.01 Viscous friction coefficient of the cart (Nsm-1) c b 0.1 Gravity (ms-2) g 9.81 Let us define the nonlinearities as follows 1 2 sin cos ; z m L z m gL (33) and consider sinmL as an uncertain term. Assuming the following bounds: 0 , 0 , 10u , and denoting max the maximal value of ; 1 max 2 max , z z and 1 min 2 min , z z the maximal and minimal values of 1 2 , z z on the domain described by the previous inequalities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002483_itec-india.2015.7386862-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002483_itec-india.2015.7386862-Figure5-1.png", + "caption": "Figure 5. General Sketch of the proposed Tracked Combat Vehicle", + "texts": [], + "surrounding_texts": [ + "The vehicle was assumed to be a series hybrid (essential in removing the bulky conventional transmission), with multiple power generators be diesel powered generator. The requirement for an energy storage capability is limited to silent watch/mobility requirements, as not much energy would be generated through braking due to the large rolling resistance associated with tracked vehicles. Power density of the powertrain is of prime importance, as more volume under armor means that more armor is required for the same amount of protection, increasing vehicle mass. Tracked drive systems can be either single line (retaining a mechanical cross-shaft) or two line systems. Twoline systems give the advantage of easier access to the rear of the vehicle. Single line systems offer superior steering performance, with drives of a lesser power rating for the same performance. A single line system is assumed for the notional vehicle. It is assumed that the vehicle will have an energy storage device, which could be batteries, flywheel or ultra-capacitor. Power flow will again be managed by the vehicle management unit across a controller area network (CAN) bus. The vehicle can have an electric suspension system with electromechanically actuated struts." + ] + }, + { + "image_filename": "designv11_34_0001824_1.4033034-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001824_1.4033034-Figure7-1.png", + "caption": "Fig. 7. Kinematic oscillation based on the analytical result (35) in the case of \u03b2 < 0. In this case, the frequency decreases with the increasing amplitude of the oscillation.", + "texts": [ + " By considering the symmetries described in the previous subsection, the lateral displacement can be expressed in the form y(\u03c8,\u03d1) = c01\u03d1+ c03\u03d1 3 + c21\u03d1\u03c8 2 +O(5), (33) where the coefficients c01,c03 and c21 are determined by (14), and the coefficient c01 gives the linear approximation of the ratio between the amplitudes of y and \u03d1. Moreover, the connection between the amplitudes of \u03c8 and \u03d1 are determined by (24), thus, we get y\u0304(\u03c8\u0304) = c01 \u221a \u2212a01 b10 \u03c8\u0304+O(3), (34) where y\u0304 is the amplitude of the oscillation expressed by the lateral displacement. By substituting (34) into (32), we get \u03c9N(y\u0304) = \u03c9L \u00b7 ( 1+\u03b2y\u03042 +O(4) ) , (35) where \u03b2 = \u2212b10 a01c2 01 \u03b2\u0302 (36) is the nonlinearity factor corresponding to the lateral displacement y. The nonlinearity factor \u03b2 expresses the increasing (Fig. 6) or decreasing (Fig. 7) tendency of the angular frequency \u03c9N if the amplitude of the oscillation is increased. The formula (35) shows Acc pt ed M an us cr ip t N ot C op ye di t Downloaded From: http://computationalnonlinear.asmedigitalcollection.asme.org/ on 03/21/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use that in case of the weakly nonlinear approximation, the difference \u03c9N(y\u0304)\u2212\u03c9L \u2248 \u03b2\u03c9Ly\u03042 between the nonlinear and linear angular frequencies is proportional to the square of the amplitude y\u0304 of the oscillation. These results can be validated by Schwab and Meijaard [18], where a special finite element method is used for the multibody simulation of the wheel-rail contact. For some typical profiles, the authors of [18] obtain results similar to those in Fig. 6. Moreover, the formula (35) can be fitted nicely to their results: in case of Fig. 7 in [18] we get \u03b2 = 0.12 \u00b71/mm2, while in case of Fig. 8 in [18], we obtain \u03b2 = 0.0023 \u00b7 1/mm2. Therefore, the frequency of first case is much more sensitive to the amplitude of the oscillation. In this section, we calculate the nonlinearity factor \u03b2 for different types of profiles that were introduced in Subsections 2.1.1-2.1.3. In the case of conical wheelset and square rails, we use the assumptions of 2.1.1, and thus the surfaces (5) and (6) become f\u00b1r (vr) = vr \u00b1b \u2212r , (37) f\u00b1w (uw,vw) = (\u2212r+huw)sinvw \u00b1(b+uw) (\u2212r+huw)cosvw " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure5-1.png", + "caption": "Fig. 5. The MacPherson strut.", + "texts": [ + " The components are as follows: The bump has been designed as shown in Fig. 4. It consists of three pairs of mounts - two of which are placed at either ends for mounting strut and one in the middle for mounting the connecting rod of slider-crank mechanism. The bump is 70 cm long and 20 cm wide and has a semicircular cross section. The mounts will be made of 3 mm thick IS2062 Grade E250 sheets. A MacPherson strut has been designed which comprises a damper and a spring. It has got proper mounting arrangements at the top and in the bottom, as shown in Fig. 5. It has been tightened with the bump and bottom platform using M12 bolts. The connecting rod is the link that connects the bump to the crank. It has holes on both ends for mounting it on the bump and on the disc crank using M12 bolts. The dimensions proposed have been used to calculate the critical buckling load Fcr in Sec- Please cite this article as: A. Sinha, S. Mittal, A. Jakhmola et al., Green energy ge ings, https://doi.org/10.1016/j.matpr.2020.06.248 tion 2.2. It will be made of 5 mm thick IS2062 Grade E250 sheets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure8.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure8.3-1.png", + "caption": "Fig. 8.3 Left: Swedish wheel, also known as Mecanum wheel, invented by the Swedish engineer Bengt Ilon. Right: Omni-wheel, an upgraded design of a omnidirectional wheel", + "texts": [ + " 198 xx List of Figures Fig. 5.6 Same cylinder with three different local reference frame assignments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243 Fig. 5.7 Gravity effect over an free-floating rigid mass . . . . . . . . . . . . . . 262 Fig. 6.1 Linear and angular velocities associated to geometric origin g of frame R1 and linear and angular velocities of the center of mass cm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 278 Fig. 8.3 Left: Swedish wheel, also known asMecanum wheel, invented by the Swedish engineer Bengt Ilon. Right: Omni-wheel, an upgraded design of a omnidirectional wheel . . . . . . . . . . . . . . . 337 Fig. 8.4 Sketch of the virtual reference frames position at the contact points of each of the three omni-wheels of an omnidirectional robot . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 337 Fig. 8.5 Differential robot Pioneer 3DX from Adept Technology, Inc. for in-door purposes, [Photograph taken from http://www", + "46) with inverse given by (7.31). The Coriolis wrench (\u2212 T (\u03bd)M\u03bd) has the form (6.70), while the gravity wrench, consider that the non-inertial frame is placed with the z-axis pointing upwards and the roll-pitch-yaw attitude representation (R-P-Y ) is chosen, as in Example6.2. Then, the gravity vector is g0 = \u239b \u239d 0 0 \u2212g \u239e \u23a0 (8.17) and the gravity wrench (\u2212MG(1)) is expressed by (6.103). 336 8 Model Reduction Under Motion Constraint The Exogenous Forces F Consider that this robot has three omnidirectional wheels, like in Fig. 8.3-right. For simplification purposes, consider the wheels to be mass-less bodies and the contact point between the floor and the wheel to be constant relative to the wheel position (refer to Fig. 8.2). For kinematic analysis, consider the sketch of Fig. 8.4, where the non-inertial frame is placed at the geometric center of the wheels such that the x-axis is positive to the front of the vehicle, the z-axis is positive in the upward direction and the y-axis is placed according to the right-hand rule" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003261_tii.2020.3004389-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003261_tii.2020.3004389-Figure5-1.png", + "caption": "Fig. 5 The rotor effect to the active winding losses (idling).", + "texts": [ + " 2 4 l l l N d w d = (3) in which, l N is the number of conductors in the l-th layer. For each conductor in the l-th winding layer, the ac loss factor ac K can be expressed as 2 4 1 0 1 1 4 l ac l l NI d K N I \u2212 = + (4) in which, 0 I is the amplitude of input current, is the skin depth at the electrical angular speed 0 2 = (5) The winding losses caused by the rotation of the rotor are calculated using full-scale FE models. The material coefficients are listed in TABLE II. It shows the rotor effects are negligible small except in the upper layers (Fig. 5) [21, 22]. The additional losses of the upper two layers are linearized as model input. The total theoretical estimated ac loss is compared with the FEAs results in Fig. 6, with satisfying result received. TABLE II MATERIAL COEFFICIENTS OF THE PMSM Core packs thickness 0.35mm Copper conductivity 57.88 x 106(\u03a9 m)-1 Iron conductivity 2.22 x 106(\u03a9 m)-1 Laminated core packs Si-Fe (V300-35A) Permanent magnet Nd-Fe-B (N42UH) Hysteresis coefficient 197 (Ws/T2/m3) Excess loss coefficient 0.16 W/(Ts -1)3/2/m3 The FEAs result shows the ac effects in the end region are much smaller than that of the active part due to no amplification effect of the iron core on the leakage magnetic field" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000333_j.ifacol.2015.06.008-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000333_j.ifacol.2015.06.008-Figure1-1.png", + "caption": "Fig. 1. Generic UVMS graph", + "texts": [ + " Choosing a single chain representation of the UVMS is advantageous for producing a general model of the system: if any part of the system is changed, the equations do not have to be rewritten, they are recomputed automatically. Furthermore, the chain representation is useful for highlighting the interaction effects between subsystems. This chain representation, also called a \u201dtree-representation\u201d consists of nodes (the links) and arcs (the joints) of the system. The reference frame for each part of the system is placed at the center of mass. In Fig. 1 a sketch of the system is presented. The vehicle is considered as a virtual 6-DOF joint, having the three translational movements of the real vehicle (surge, sway, heave) and the three rotational movements (roll, pitch, yaw). For simplification of the kinematic problem, the virtual 6-DOF joint can be decomposed into six independent joints with zero mass. The connection between the first joint of the manipulator and last DOF of the vehicle is represented through a node that preserves the characteristics of the vehicle: mass, length, radius" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000693_gt2014-25904-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000693_gt2014-25904-Figure1-1.png", + "caption": "Figure 1. PHOTOGRAPH OF THE TEST RIG.", + "texts": [ + " A number of experiments on an experimental rig with bladed discs were conducted using 2 types of 8-bladed discs of (long and short blades) with healthy but mistuned, crack in one blade and a crack in two blades, and data are recorded into a PC during the machine run-up using the shaft torsional vibration measured using the incremental shaft Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 10/16/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2014 by ASME encoder. The paper presents the experimental setup, simulation of blade faults, experiments conducted, observations and comparison of results between the long and short blades with and without faults to reveal which of these types is more effective for the purpose of BHM. An experimental test rig is shown in Figure 1 [1] which is located in the Dynamics Lab at the School of MACE in the University of Manchester. The test rig mainly consists of 1 HP driver 3-phase motor, speed controller for run up, cost down or steady state speed operation, 2 ball bearings, a 20 mm diameter steel shaft with a bladed disc. The shaft is connected to the motor through a flexible coupling. The rig photograph is shown in Figure 1 which includes instrumentation as well. The measuring system is composed of one rotary shaft encoder for the shaft IAS, accelerometers on bearings and a tacho sensor used to monitor rotating speed [1]. The whole schematic of the data acquisition system is shown in Figure 2. The bladed disc consists of 8 rectangular blades. Typical photographs of the 2 bladed discs (long and short blades) used in the study are shown in Figure 3. The photographs of both long and short blades with measuring scale are also shown in Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002205_0959651816655039-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002205_0959651816655039-Figure4-1.png", + "caption": "Figure 4. 3-DOF hover helicopter.", + "texts": [ + " The dynamic model of the attitude system of the quadrotor can be expressed as \u20acui = ui1 + (Jf Jc) Ju _ci _fi \u20acfi = ui2 + (Jc Ju) Jf _ci _ui \u20acci = ui3 + (Ju Jf) Jc _fi _ui 8>>>>< >>>>>: \u00f024\u00de where i=1, 2, . . . , n. For the ith quadrotor, the states consist of ui, _ui, fi, _fi, ci, _ci which represent the pitch angle and rate, the roll angle and rate, the yaw angle and rate, respectively. The three inputs ui1, ui2, ui3 are the angular accelerations applied in the u, f, c directions. Ju, Jf, Jc represent rotational inertia of the pitch axis, roll axis and yaw axis, respectively. Let us take the 3-DOF hover helicopter shown in Figure 4, which is produced by Quanser Company as the research object to simulate the operation of the quadrotor attitude system. The 3-DOF hover helicopter mainly consists of electric motors, rotors, helicopter body, power-supply module and encoders (sensors). In the existing software platform of the 3-DOF hover, we can design the proposed FCCT system under the circumstance of MATLAB REAL-TIME. With the help of Quansers supporting software, the block diagrams of MATLAB Simulink can be directly encoded into C language which downloads to the real-time simulation system from a supporting PIC card through the parallel port" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001644_ijrapidm.2015.073548-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001644_ijrapidm.2015.073548-Figure5-1.png", + "caption": "Figure 5 Surface comparison PGAM: sample 3", + "texts": [], + "surrounding_texts": [ + "Table 3 shows the overall linear and angular dimensions of the produced sample. The figure of standard uncertainty is also presented per dimension and per machine. Consistent with the results displayed in Table 3, all the dimensional features in the Z-axis (i.e. dimensions related with the bottom plane of the produced parts) have shown a much higher deviation. In this regard, the average error value of dimension 6 in the CGAM produced parts has grown to 2.245 mm and is not displayed in the figure. Figure 4 shows the average linear dimensional error of the parts after the hybrid manufacturing process. The errors have been calculated by computing the difference between the average dimensions of the produced sample and the nominal value. Results show that the dimensional errors related with the bottom plane of the part (i.e. attaching plane on the additive machine) suffer from much higher deviations due to the warping effect typically present in material extrusion processes. The results displayed in Figure 4 and Table 3 demonstrate that linear dimensions and angular dimensions on the horizontal plane (XY plane) have very low deviations and the performance is similar in both machines. According to the general tolerance standard (ISO, 2000), the tolerance class designation of the PGAM sample has fallen under the category of \u2018fine\u2019. In the case of the CGAM sample, average results fall under the same tolerance category. However, standard uncertainty is higher and certain dimensions, such as numbers 9 and 11, show poor results falling under the category of \u2018medium\u2019. Figures 5 and 6 show the CAD to part surface comparison of PGAM sample 3 and CGAM sample 2, respectively. Red areas represent the dimensions which are above the nominal CAD model and blue areas represent dimensions which are below the nominal CAD model. Deviation to nominal CAD model is in the range of \u20130.40 mm to 0.40 mm for PGAM produced parts and \u20132 mm to 2 mm for CGAM produced parts. At the same time, the bottom view of the figures shows that warping is much more severe in CGAM produced parts, thus having a deviation maximum up to 3.4 mm. Consistently, with the trends previously displayed in Figure 4 and Table 3, the top view of the visualised samples demonstrates that dimensional deviations are much lower in all the post-processed surfaces (i.e. XY plane), hence subtractive processes are increasing the dimensional stability. Figure 7 shows the General Dimensions and Tolerances (GD&T) studied during this experiment, and the data displayed correspond to sample 3, manufactured with CGAM equipment. Average DG&T values and corrected standard deviations for small samples are shown in Table 4. The cylindricity and roundness tolerances in holes 1 and 2 (GD&T 1, 2, 3 and 4) show equivalent results for CGAM and PGAM parts. The top plane and front plane perpendicularity (GD&T 5), the front plane and middle plane parallelism (GD&T 6) and the top plane flatness (GD&T 7) show better results for PGAM parts. The majority of these GD&T features fall under the tolerance class K, according to standards, not depending on the type of machine used. Table 4 also shows that thermal warping had a negative effect on GD&T features, such as the bottom plane and top plane parallelism (GD&T 8), the bottom plane and front plane perpendicularity (GD&T 9) and the bottom plane flatness (GD&T 10). They have been computed using the best fit plane of the bottom plane of the produced sample, and therefore its deviation is much higher. The qualitative evaluation of the tested additive machines showed that the impact of initial layers over the overall geometry is fundamental when using CGAM equipment. Figure 8 shows how the raft of the CGAM system has been detached from the base plane after one hour of manufacturing due to the thermal warping, which creates stress that causes the detachment of the geometry from the base plate. The CGAM system failed several times during this experiment, and the manufacturing process was unpredictable. In addition, as it is shown in Figure 9, the filament of the CGAM produced part had bad cohesion for the subtractive process. All the CGAM produced parts showed delamination of the filament in contact with the milling cutter, which also affected the final geometry. Nevertheless, dimensional stability was substantially improved and, for instance, the stair-step effect was eliminated after the milling process." + ] + }, + { + "image_filename": "designv11_34_0001035_09205071.2014.963698-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001035_09205071.2014.963698-Figure2-1.png", + "caption": "Figure 2. Plan view of the circular rings with parallel axes.", + "texts": [ + " To fulfill this requirement, we replace the exponential factor of (1) with the ansatz [11], which separates the total field into the source and secondary parts: Ae zt \u00fe Bezt \u00fe e z f1j jt (3) Then, the boundary condition Br l\u00f0 \u00de \u00bc 0 leads to Ae lt \u00fe Belt e l f1\u00f0 \u00det \u00bc 0 Aelt \u00fe Be lt \u00fe e l\u00fef1\u00f0 \u00det \u00bc 0 (4) Solving (4) gives A \u00bc cosh l f1\u00f0 \u00det\u00bd cosh lt\u00f0 \u00de sinh lt\u00f0 \u00de\u00bd sinh 2lt\u00f0 \u00de B \u00bc cosh l \u00fe f1\u00f0 \u00det\u00bd cosh lt\u00f0 \u00de sinh lt\u00f0 \u00de\u00bd sinh 2lt\u00f0 \u00de 8>< >>: (5) And hence, we have the vector potential of the circular ring shielded by two parallel screens: A l\u00f0 \u00de u r; z\u00f0 \u00de \u00bc l0I1r1 R1 0 cosh l \u00fe f1\u00f0 \u00det\u00bd cosh l z\u00f0 \u00det\u00bd sinh 2lt\u00f0 \u00de J1 rt\u00f0 \u00deJ1 r1t\u00f0 \u00dedt; l f1\\z l (6) Providing that there is another circular ring with the radius r2, located in the plane z = \u03b62 ( l\\f1\\f2\\l) and carrying the electric current I2, and the distance between the parallel axes of both rings is \u03c10, and r \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi q20 \u00fe r22 2q0r2 cosw p (See Figure 2). Then, the mutual magnetic energy is given by W mutual\u00f0 \u00de m \u00bc 1 2M12I1I2 \u00fe 1 2M21I1I2 \u00bc MI1I2 \u00bc I2r2 R2p 0 A l\u00f0 \u00de u r; f2\u00f0 \u00de cos p w\u00fe u\u00f0 \u00de\u00bd dw \u00bc I2r2 R2p 0 A l\u00f0 \u00de u r; f2\u00f0 \u00de cos w\u00fe u\u00f0 \u00dedw \u00bc 2I1I2l0pr1r2 R1 0 cosh l \u00fe f1\u00f0 \u00det\u00bd cosh l f2\u00f0 \u00det\u00bd sinh 2lt\u00f0 \u00de J1 r1t\u00f0 \u00deJ1 r2t\u00f0 \u00deJ0 q0t\u00f0 \u00dedt (7) Finally we obtain the mutual inductance between two circular rings shielded by two parallel screens of infinite permeability: M \u00bc 2l0pr1r2 Z1 0 cosh l \u00fe f1\u00f0 \u00det\u00bd cosh l f2\u00f0 \u00det\u00bd sinh 2lt\u00f0 \u00de J1 r1t\u00f0 \u00deJ1 r2t\u00f0 \u00deJ0 q0t\u00f0 \u00dedt (8) For l\\f2\\f1\\l, \u03b62 and \u03b61 of (8) should be exchanged" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001776_iemdc.2015.7409253-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001776_iemdc.2015.7409253-Figure2-1.png", + "caption": "Fig. 2. FEM model of induction motor a) Healthy b) Broken bars fault", + "texts": [ + " I \u2013 Parameters of the test motor Number of poles 4 Outer diameter of stator, mm 145 Inner diameter of stator, mm 89 Outer diameter of rotor, mm 88 Inner diameter of rotor, mm 34.7 Number of stator slots, mm 36 Number of rotor bars 28 Rated voltage, V 400 Rated frequency, Hz 50 Rated power, kW 2.2 Rated current, A 4.7 Rotor inertia, kgm2 0.047 Fig. 3. Stator phase currents a) Healthy b) Motor with two broken bars parameters of the test motor are used in the simulations. Electrical and mechanical parameters of the motor are given in Table. I. The simulation model shown in Fig.2 is designed based on the actual size and geometry of all parts, such as stator core, rotor core, shaft, rotor bar structure. Furthermore, stator winding configuration, air gap, physical conditions of the stator conductors, stator and rotor magnetic material are also taken into account. For broken bar condition, the corresponding rotor bars are deleted from the model to block the bar currents. Fig. 3 shows one phase stator current in case of healthy and two broken bars condition. It is clear in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001123_amm.660.633-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001123_amm.660.633-Figure5-1.png", + "caption": "Fig. 5 Deformation behaviour of superstructure bus frame in quasi-static test", + "texts": [ + " The simulation results indicate a variation of negative plastic strain around the frame at the early stage of impact. The non-linear trend was consistent due to unstable distribution of impact. The rollover test simulation shows that R-side of structure deform towards the concrete surface, L-side deform towards the impact point, the roof structure deform outwards, away from bus occupant as shown in Fig. 4. The deformation behaviour of frame structure undergoing quasi-static loading test is depicted by Fig. 5. It can be observed that R-side deform towards the inner compartment, L-side deforms away from occupant space and the roof deformation follows the direction of impact from right to left. Fig. 6 shows that the maximum stress experienced by the middle frame of R-side structure, while the back compartment hit the maximum total deformation as shown in Fig. 7. During quasi-static load impact, the maximum stress measured around a middle structure of the bus frame, and the maximum total deformation occurs on the side impact of front compartment as shown in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000385_9781118682975.ch8-Figure8.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000385_9781118682975.ch8-Figure8.6-1.png", + "caption": "Figure 8.6 Manipulation of hydration number to modulate relaxivity of metal ion-responsive MRI contrast agents", + "texts": [ + " The desire to visualize biological processes by MRI has sparked a trend in the development of next generation MRI contrast agents that can be biologically activated by a specific event. By optimizing the three parameters of relaxivity (q, \ud835\udf0fm, and \ud835\udf0fR), these \u201csmart\u201d contrast agents can selectively undergo relaxivity change in response to a specific biomarker or analyte of interest, for example, metal ions. Manipulation of the hydration number q is the most common strategy for development of metal ion-activated MRI contrast agents. According to relaxation theory, enhanced water accessibility to the paramagnetic metal center yields higher relaxivity. As illustrated in Fig. 8.6, the Gd3+-based contrast agents contain a paramagnetic Gd3+ chelate core and at least one metal ion receptor. In the apo form of the Gd3+ complex, an arm of the metal receptor binds the Gd3+ center and blocks water access. In the presence of the sensed metal ion, this metal receptor dissociates from the Gd3+ center in favor of binding the sensed metal. This intramolecular ligand displacement reaction opens an additional binding site for one water molecule to reach the inner sphere of Gd3+, changing q from 0 to a higher level (q = 1) upon metalation and providing enhanced signal contrast" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000788_metroaerospace.2014.6865903-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000788_metroaerospace.2014.6865903-Figure6-1.png", + "caption": "Fig. 6. Voxel\u2019s tandard deviation in a transversal section of the the acquisition.", + "texts": [ + "5 \u03bcm (obtained with test comparable to EN/ISO 10360-2) and can acquire approximately 70,000 points/s, and 900 points per line. The CMM system has a range of accuracy of about 2\u00f76 mm with a positioning maximum speed of 750 mm/s. To avoid reflections the surface has been white-mat painted. Doing so, after a multiple view recomposition, a cloud of 479780 points has been achieved and analysed according to a voxel structure of 64x64x64 voxels, as shown in Fig. 5. Following the segmentation scheme described in section 2, planar surfaces and cylinders are recognized. Fig. 6 shows the voxel\u2019s values of the standard deviations from the local best fitting plane, taken in a transversal section of the flange. Fig. 6 shows that planar surfaces always have standard deviation smaller then the cilindric ones, with a ratio of about 1:2. This allows to find a significant threshold for distinguishing cylinders from planes. The IRS is found starting from the direction orthogonal to the transversal section of Fig. 6. The second axis is found as the most populated set of aggregated voxels with a direction normal to the first one. It results to be one of the set of the five protrusions at 72\u00b0, as shown in the blueprint of Fig.4. The third axis derives from vectorial product. From IRS the recognition of cylindric voxels is found through the osculating radius evaluation. Fig. 7 shows the map of the voxel\u2019s osculating radii of the sections with smaller holes, confirming a good resolution of the proposed method. For example considering the 4 smaller holes in the left part of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003427_2050-7038.12578-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003427_2050-7038.12578-Figure1-1.png", + "caption": "FIGURE 1 The cross section of a 8/6 SRG with flux lines: A, aligned position, B, unaligned position", + "texts": [], + "surrounding_texts": [ + "As rotor and stator of the switched reluctance machine have salient poles, the phase inductance of the machine is nonlinear and it varies from maximum to minimum as the rotor moves from the aligned to unaligned position with respect to the active stator pole shown in Figures 1 and 2. Magnetic energy stored in the machine tends to retain the aligned position for the rotor. While the prime mover forces the rotor to pass the aligned position, a negative electromagnetic torque will be applied to the rotor which results in production of the back-EMF in the machine and generating electrical power. As there is no permanent magnet and windings on the rotor, an excitation current pulse in the aligned position should be applied to the stator phase winding. To prepare the required magnetic field in the machine for power generation, it must be noted that current direction in the winding is of no importance in production of the electromagnetic torque.26 After building the magnetic field, mechanical energy of the prime mover can be converted into electrical energy. This energy is transferred through the converter. The phase voltage equation in the generating mode is as follows: where Vph, e, R, i, L, \u03c9, and \u03b8 are phase voltage, back-EMF, winding resistance, phase current, phase inductance, angular speed and rotor position, respectively. The value of back-EMF is a function of current, rotor speed and the first derivative of the phase inductance. Three types of current waveform are possible for the generator according to the value of back-EMF which is depicted in Figure 3. Regarding the parameters defined in this figure, \u03b8on is turn-on angle, \u03b8off is turn-off angle and \u03b8ext is the position in which the current reaches to zero. During the period of excitation (between \u03b8on and \u03b8off), the power is supplied from the dc source to the generator and after turn-off angle, the generated current flows back to the dc source or the load. With regard to (1), when the backEMF is greater than the summation of applied voltage and the resistive voltage drop, the current will increase even after applying the negative voltage to the phase as illustrated for the first type of phase current waveform depicted in Figure 3. When the back-EMF and the sum of the applied voltage and the resistive drop are in balance, the current waveform will be flat-top as observed in the second type of phase current waveform shown in Figure 3. Finally, when the value of negative applied voltage during generating mode is greater than back-EMF, the current will fall back to zero as illustrated for the third type of phase current waveform depicted Figure 3. For the wind generation system, the operating speed of the generator is in the low and medium speed ranges. Also, maximum output power is obtained in flat-top current waveform mode.25 When the generator is used in the wind energy system, the speed and current set points are determined based on the interaction with the wind speed to capture the maximum possible energy from the wind turbine. So, the value of back-EMF is completely dependent on the wind velocity. As a result, the dc bus voltage FIGURE 2 Idealized phase inductance profile1 should be tuned to be in balance with the back-EMF to obtain the maximum output power. In the following section, a novel control method is presented to achieve the flat-top current waveform for the generator." + ] + }, + { + "image_filename": "designv11_34_0002246_978-3-319-15684-2-Figure13.7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002246_978-3-319-15684-2-Figure13.7-1.png", + "caption": "Fig. 13.7 3D model of a part manufactured by milling (courtesy OAO TulaTochMash, Tula, Russia)", + "texts": [ + " We have deducted the following generalized regression equations that express the relations between machining time T, surface area F, and cutting feed Sp [6]: T Sp; F\u00f0 \u00de \u00bc 3x3 \u00fe 3y3 \u00fe 3z3 \u00fe 4:304x 0:061y; T Sp; t\u00f0 \u00de \u00bc 3x3 \u00fe 3y3 \u00fe 3z3 0:121y2 \u00fe 1:497z2 6:422xy\u00fe 0:016yz\u00fe 12:806x 0:062; T F; t\u00f0 \u00de \u00bc 3x3 \u00fe 3y3 \u00fe 3z3 \u00fe 0:979x 4:695y\u00fe 3:452z 0:519: The graphs for these equations are shown in Fig. 13.5. The following regression equations have been produced for primary milling features (Table 1): 112 D.I. Troitsky 13 Conflicts in Product Development \u2026 113 Machining Time Model Verification To validate the proposed model we will apply it to estimate the machining time for milling the part shown in Figs. 13.6 and 13.7. The part has 16 surfaces to be machined (see Fig. 13.7): two holes, three slotted surfaces, and nine planes. They are listed in the Table 2: The total machining time is a sum of machining times for the holes, the slot, and the other surfaces. Surfaces 1 and 10, 2 and 11 are jointly machined. 2 holes 114 D.I. Troitsky 13 Conflicts in Product Development \u2026 115 The machining time estimated with the available reference tables is 5.2 min. The difference is about 11 %, so the proposed model is valid and feasible. The proposed machining time estimation model can be used to avoid professional conflicts in product development" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001901_110505-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001901_110505-Figure4-1.png", + "caption": "Fig. 4. (color online) The states of agents with \u03c4s = 0.1, T = 0.2, and \u03b1 = 1.", + "texts": [ + " 110505-7 When the second-order multi-agent systems (4) apply the consensus tracking protocol (6), each agent can track timevarying reference state with \u03c4s = 0.1,T = 0.3,\u03b1 = 1 in Fig. 2. Figure 3 shows the divergence curves under the conditions of \u03c4s = 0.5, T = 1, \u03b1 = 1, which means that the system cannot achieve the consensus tracking. Consider a special case, if the given reference state of virtual leader is (x0(kT ) ,y0(kT )) = (3,0), then the consensus tracking problem is simplified to the aggregating problem. From Fig. 4, it is easy to see that when \u03c4s = 0.1,T = 0.2, \u03b1 = 1, which satisfy the given conditions in Corollary 2, the consensus position states of agents 1\u20135 are three, and the speed states are zero. However, in the case of the multi-agent system (4), applying the consensus tracking protocol (6) with m = 0 and \u03c4s = 0.5, T = 1.2, and \u03b1 = 1, cannot achieve the aggregating is shown in Fig. 5. Then based on the above numerical results, the results of Theorem 1 are numerically verified. 6. Conclusions The bounded consensus tracking problems of secondorder multi-agent systems based on sampled-data with sampling delay under directed networks have been studied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure8-1.png", + "caption": "Fig. 8. The designed gear train.", + "texts": [ + " A disc shaped crank has been used because it is easy to design and manufacture. It will be made of 5 mm thick IS2062 Grade E250 sheets. A compound gear train has been designed and assembled to serve as the power transmission and speed multiplication unit. Spur gears have been used with specifications and dimensions as mentioned in Section 2.3. Gears have been meshed by aligning them with their pitch circles that are in turn, tangential to each other. Loose gear as shown, will be connected to the generator shaft. Fig. 8 shows the assembled design of the same. The complete assembly of the model has been done containing all the components as listed above, including bearings and bearing neration from road traffic using speed breakers, Materials Today: Proceed- The Finite Element Analysis (FEA) was done on Ansys 16.0 Student version which is a Computer Aided Engineering (CAE) software. Static structural analysis has been done for both the below components. Only these components were chosen for FEA as they are the most critical components having high chances of failure at various points in them" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000753_amm.393.403-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000753_amm.393.403-Figure4-1.png", + "caption": "Fig. 4: Meshed Gear Pair Model with boundary conditions", + "texts": [ + " The 2D CAD model of the spur gears was built for the contact pressure analysis using FEA. The solid model was then imported into ANSYS software version 12.0 through IGES format and meshed accordingly. Gear pair is meshed using Plane 42 element [11], a 2D element type. The regions near the point of contact and near fillet are meshed fine in order to achieve a good accuracy in the results. The contact between the gears was modelled using CONTA 175 and TARGE 169 elements [11]. The boundary conditions are applied in the solution stage as shown in Fig. 4. These conditions and loading are changed accordingly for different rotation angle of gear pair. The load distribution along the tooth flank is done according to Fig.1. FEM model is prepared such that it suitably matches with the twin-disc test conducted in the previous work. The FEA results for the contact pressure distribution at different positions along the contact path are shown in Fig. 5. The FEA results of the gear pair along the line of action are listed in Table 2 together with analytically calculated Hertzian pressure and twin disc test results obtained from the literature [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.3-1.png", + "caption": "Figure 9.3 Left and right shouldered poses of a Puma manipulator.", + "texts": [ + "57) If s\ud835\udf035 \u2260 0, the other equation pairs give \ud835\udf034 and \ud835\udf036 as follows without additional superfluous sign variables: \ud835\udf034 = atan2(\ud835\udf0e5c\u221723, \ud835\udf0e5c\u221713) (9.58) \ud835\udf036 = atan2(\ud835\udf0e5c\u221732,\u2212\ud835\udf0e5c\u221731) (9.59) (a) First Kind of Multiplicity The first kind of multiplicity is associated with the sign variable \ud835\udf0e1 that arises in the process of finding \ud835\udf031 by using Eqs. (9.39) and (9.40). The manipulator attains the same location of the wrist point by assuming one of the two alternative poses that correspond to \ud835\udf0e1 = + 1 and \ud835\udf0e1 = \u2212 1. These poses are illustrated in Figure 9.3. The poses corresponding to \ud835\udf0e1 = + 1 and \ud835\udf0e1 = \u2212 1 are designated, respectively, as left shouldered pose and right shouldered pose. In both poses, as implied by Eq. (9.34), the angle \ud835\udf191 denotes the projected direction of the line OR on the 1\u20132 plane of the base frame. Meanwhile, Eq. (9.39) implies that the angle \ud835\udf131 becomes acute (i.e. \ud835\udf131 <\ud835\udf0b/2 with c\ud835\udf131 > 0) if \ud835\udf0e1 = + 1 and it becomes obtuse (i.e. \ud835\udf131 >\ud835\udf0b/2 with c\ud835\udf131 < 0) if \ud835\udf0e1 = \u2212 1. On the other hand, \ud835\udf031 = \ud835\udf191 \u2212\ud835\udf131 according to Eq. (9.40). That is why \ud835\udf031 assumes the positive and negative values shown in Figure 9.3 when the manipulator is in one of the left and right shouldered poses. Of course, \ud835\udf032 also assumes a value that is compatible with one of these poses as implied by Eqs. (9.50) and (9.51). (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e3 that arises in the process of finding \ud835\udf033 by using Eqs. (9.44)\u2013(9.46). The manipulator attains the same location of the wrist point by assuming one of the two alternative poses that correspond to \ud835\udf0e3 = + 1 and \ud835\udf0e3 = \u2212 1", + "52) that was derived for the Puma manipulator. Therefore, it can be solved similarly for the wrist joint variables as explained in Section 9.1.3. (a) First Kind of Multiplicity The first kind of multiplicity is associated with the sign variable \ud835\udf0e1 that arises in the process of finding \ud835\udf031 by using Eq. (9.142). The manipulator attains the same location of the wrist point by assuming one of the two alternative poses that correspond to \ud835\udf0e1 = + 1 and \ud835\udf0e1 = \u2212 1. These poses are very similar to those illustrated in Figure 9.3 for the Puma manipulator. Therefore, they are also designated as left shouldered if \ud835\udf0e1 = + 1 and right shouldered if \ud835\udf0e1 = \u2212 1. They are illustrated in Figure 9.14. (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the wrist joint variables (\ud835\udf035, \ud835\udf034, \ud835\udf036). Therefore, it is the same as the third kind of multiplicity of the Puma manipulator. In other words, as illustrated in Figure 9.5, it occurs similarly as a wrist flip phenomenon without any visual distinction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001802_978-3-662-46312-3_29-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001802_978-3-662-46312-3_29-Figure1-1.png", + "caption": "Fig. 1 Principle tool concept for the fine blanking of helical gears", + "texts": [ + " The decisive advantage of fine blanking is the achievable high quality of the blanked part. To reach this high quality in rotary fine blanking it is necessary to ensure a small die clearance. The typical die clearance is approximately 0.5 % of the sheet thickness. The interaction of the die clearance, v-ring and the superposition of compressive stresses are responsible for the achieved part quality [5]. To realize the fine blanking of helical gears tool kinematics had to be redesigned. The linear movement of the active parts gets superimposed with a rotational movement (Fig. 1). Thereby, the punch and counter punch are mounted pivoted and axially fixed. The upward movement of the cutting plate causes the punch to the rotational movement that leads to a helical cut. The PhD thesis of Zimmermann gives a detailed process description for the fine blanking of helical gears made of 16MnCr5 [6]. Technological challenging was the development of the tool redesign. The rotation angle \u02c7 is only 1.4\u00b0 and thereby the realization of a pivoted bearing is difficult to implement. The use of rolling bearings was not possible due to process related limitations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000834_12.2083777-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000834_12.2083777-Figure1-1.png", + "caption": "Figure 1: (a) Schematic description of the laser micro sintering process. The powder material is transported by an automatic refillable ring rake. The powder is coated in small layers during a circular motion of the ring rake over the building piston. A focused laser beam is finally sintering the lose powder material. (b) One of the developed sinter stations, the laser, optical setup and controlling logic are fully integrated. A powder handling subsystem (left) is attached. (c) Schematic view of the main components of the sintering setup. For higher deflection velocities an additional acousto-optic modulator is used to switch the laser beam precisely enough for microscale resolution.", + "texts": [ + " The initial studies of LMS indicated that a stochastic pulse distribution results in a higher stability and repeatability of the sinter process [5]. This was observed in particular for direct absorbing material powders (e.g. metals) with a reduced heat conduction of the powder bed where the grain sizes are smaller than that length the heat can diffuse during the irradiation time. However, for non-direct absorbing materials (e.g. ceramics) the opposite effect was observed [6], [7]. Since 2001, the laser micro sinter machines were constructed with respect to the specific sinter process requirements, as schematically shown in Fig. 1: Powder coating of the sinter layer was done by a cylindrical ring rake that temporarily carries a defined amount of powder material until a refill by the included powder piston. The degree of filling and the coating velocity strongly depends on the flowability of the applied powder material which is influenced, among others, by the grain size, grain distribution, humidity, etc. Also the shape of the coating blade (at the bottom of the rake, from sharp to blunt) has to be prepared according to the applied powder material. The ring rake is guided and propelled by a lever to sweep across the powder bed in a circular motion, ideally leaving a thin powder layer that will be sintered afterwards according to the respective cross section of the intended body. The sintering setup is located in a vacuum chamber [Fig. 1(c)], which allows to work under arbitrary atmospheres. In particular by using metallic powders with particle diameters below 2 \u00b5m, the reactive surface is so enlarged that first laser irradiation will induce massive exothermic reactions. In order to prevent this undefined oxidation, the oxygen within the process chamber was exchanged by inert process gases, such as helium or argon. In special cases [7], the sinter chamber was applied to work under reactive atmospheres. Investigations of LMS under reduced pressure were also performed and showed the possibility to modify the morphology and the porosity of the sintered micro parts [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002793_6.2020-1721-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002793_6.2020-1721-Figure3-1.png", + "caption": "Fig. 3 Relationship between the (inertial) relative frame R and the body-fixed frame B.", + "texts": [ + " In the definition of the angles in Figure 2, arrows indicate direction of increasing angle and an D ow nl oa de d by U N IV E R SI T Y O F T O R O N T O o n Ja nu ar y 6, 2 02 0 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 02 0- 17 21 Fig. 1 Schematic representing a standard carrier landing approach profile. This includes the glideslope trajectory, defined by a flight path angle \u03b30 and nominal velocity V0, which is used to guide the UAV into the final landing maneuver. angle is zero when the vector is parallel to the horizontal, such that \u03b80 = \u03b10 + \u03b30. Figure 3 shows the relationship between the relative frame R and the body-fixed frame B. The coordinates \u03b4px , \u03b4pz , and \u03b4\u03b8 are with respect to the relative frame R. The dynamics of the UAV, described in Section III, will be written with respect to the relative frame. Fig. 2 Relationship between the earth-fixed frame I and the (inertial) relative frame R. Now that the relevant reference frames have been defined, we can discuss the dynamics model for the UAV system. While classical approaches to modeling aircraft dynamics typically require aerodynamic coefficients regressed from data, we use a principled CFD model of the aerodynamics that is directly coupled with the rigid-body motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002509_978-3-319-49259-9_30-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002509_978-3-319-49259-9_30-Figure2-1.png", + "caption": "Fig. 2. Worst case executions of the plane formation algorithm [10] in the SSYNC model with non-rigid movement. The existing algorithm lands the robots to the target plane shown in (a). There exists an SSYNC schedule that allows convergence to a plane (b), and non-rigid movement allows regular octahedron (c).", + "texts": [ + " The impossibility is clear because the SSYNC model with nonrigid movement allows the worst case executions in the FSYNC model with rigid movement. To show the solvability, we present a novel plane formation algorithm for the SSYNC robots with non-rigid movement, because the existing plane formation algorithm [10] relies heavily on the FSYNC schedule and rigid movement. We start with an example that shows the effect of semi-synchrony and nonrigid movement on the existing plane formation algorithm in [10]. Consider an initial configuration where six robots form a triangular anti-prism (Fig. 2(a)). This configuration satisfies the condition (i) of Theorem 1. The existing plane formation algorithm makes the robots agree on the plane that is parallel to the bases and contains the center of the triangular anti-prism, and sends each robot to the plane along the perpendicular to the plane. Consider an SSYNC schedule of the robots (with rigid movement) where in an odd time step the robots forming the top base execute a cycle, and at an even time step, the robots forming the bottom base execute a cycle. Then the robots converge to a plane, but they never land on one plane (Fig. 2(b)). Consider non-rigid movement in the FSYNC model, for example, the robots stop at the points where each side face of their anti-prism forms a regular triangle. Thus they form a regular octahedron (Fig. 2(c)). Though the algorithm of [10] guarantees that the robots can form a plane from a regular octahedron, during the formation, the algorithm allows the robots to form a triangular anti-prism. Thus the robots may fall into a loop between a regular octahedron and a triangular anti-prism.1 Additionally, the correctness proof of the go-to-center algorithm in [10] heavily relies on the rigid movement. The authors consider all possible next positions, that form a uniform polyhedron. Because the robots cannot select all the vertices of the uniform polyhedron, the go-to-center algorithm succeeds in symmetry breaking" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002229_acc.2016.7525039-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002229_acc.2016.7525039-Figure8-1.png", + "caption": "Figure 8. Logarithmic norm of the complementary sensitivity function vs. logarithmic dimensionless frequency, \u03c4\u0304\u03c9, for different values of \u03b5. \u03b3 = 0.3 and \u03bb \u03c4\u0304 = 1.5. The surfaces are plotted for three different levels of gain relative error, \u03b4 = 0,\u00b10.3. Complementary sensitivity function shows a better performance with zero steady-state error at low frequencies and smaller magnitude at higher frequencies.", + "texts": [], + "surrounding_texts": [ + "The IMC method was proposed to design a robust control system for a linear first-order time-delay plant under large bounds of uncertainties using small-gain theorem. Based on the robust stability condition an explicit criterion for tuning the controller parameter is proposed. The worst case to tune the controller parameter was shown to be the case with the lowest values for the delay and the system lag and the largest value of the system gain. Complementary sensitivity function is invoked to compensate for the system performance. The maximum value of the complementary sensitivity function is used to make a trade-off between the system robustness and the performance. Performance characteristics of the IMC was demonstrated through a setpoint tracking objective. IMC successfully achieves robust stability and in the meantime it tracks accurately the reference signal with roll-off characteristics at high frequencies." + ] + }, + { + "image_filename": "designv11_34_0002364_stab.2016.7541184-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002364_stab.2016.7541184-Figure1-1.png", + "caption": "Fig. 1. Cylinder.", + "texts": [ + " Further problem of reactions distribution R\u0303\ud835\udc56 in some fixed point of time is investigated by the proposal that force \u03a6\u0303 is acting at the point r\u0303\ud835\udc36 and force moment there is zero. Motion equations (1) for finding reactions of walking robot body and legs prescribed motion can be transformed [5]\u2013[8]: \u2211\ud835\udc5b \ud835\udc56=1 R\u0303\ud835\udc56 = \u03a6\u0303, \u2211\ud835\udc5b \ud835\udc56=1 r\u0303\ud835\udc56 \u00d7 R\u0303\ud835\udc56 = r\u0303\ud835\udc36 \u00d7 \u03a6\u0303. (2) Assuming that the robot footholds are on the surface of a rough cylinder of radius \ud835\udf0c with a friction coefficient \ud835\udc58, we introduce the coordinate system \ud835\udc42\ud835\udc65\ud835\udc66\ud835\udc67 such that the axis \ud835\udc42\ud835\udc65 is directed along the cylinder axis (so that the projection of \u03a6\u0303 on the axis \ud835\udc42\ud835\udc65 is negative \u2013 see Fig. 1.), the axis Oz is parallel to the vector \u03a6\u0303, and the angle between the cylinder axis and the vector \u03a6\u0303 is \ud835\udefc. The problem of finding the reaction forces (2) is similar to the foothold reactions distribution problem in homogeneous gravity field, when the footholds are on the external surface of a rough inclined cylinder where the axis has an angle \ud835\udefc with respect to the vertical vector \u03a6\u0303. It has been considered in Ref. [6], where the problem of searching the reactions components along the cylinder axis when \ud835\udefc = 0 and the work [8] for arbitrary \ud835\udefc was considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001507_ijca.2015.8.2.16-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001507_ijca.2015.8.2.16-Figure2-1.png", + "caption": "Figure 2. Planetary Gear Dynamic Model", + "texts": [ + " The paper is arranged as follows: Section 2 establishes an elaborate dynamic model. A novel processing method in Hamiltonian system is proposed in Section 3. Section 4 carries out simulation to verify the method. Conclusions are given in Section 5. load through a 2K-H planetary gear reducer. A planetary gear dynamic model is consist of the sun gear (subscripts s ), the planet gear (subscripts pi , 1,2i n ), the ring gear and carrier (subscripts c ), where each stage has n planet gears as shown in Figure 2. Copyright \u24d2 2015 SERSC 155 The nonlinear model is established by the lumped parameter method in the following hypotheses: (1) All the planetary gears are assumed to be spur gears with the same physical and geometrical parameters. (2) The influence of friction in the process of the gear mesh can be neglected. Each corresponding mathematical model (Figure 1) is obtained by Newton-Euler method as follows: The analysis of the sun gear: 1 [ ] n s s D spi spi s i J T P D r (1) The analysis of the ith planet gear: +pi pi spi spi rpi rpi piJ P D P D r (2) The analysis of the carrier: 2 1 1 ( ) + + cos( ) n n c pi c c spi spi rpi rpi c L i i J m r P D P D r T (3) Where, J (subscripts s , pi , c ) is the inertia ranged from sun gear to planet carrier, sr is the base circle radius, cr is the carrier radius which is equivalent to the pitch circle radius of sun gear add to the pitch circle radius of planet, m (subscript pi ) is mass of planet gear, is the mesh angle of gears, DT is the input matrix, LT is the load matrix, and P , D (subscripts spi , rpi ) are elastic mesh force and viscous mesh force, in which subscript spi represents the internal meshing between sun gear and the planet gear and subscript rpi represents the external meshing between the planet gear and carrier, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003789_j.sna.2020.112386-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003789_j.sna.2020.112386-Figure2-1.png", + "caption": "Fig. 2. Schematic and cross-section of a liquid crystal micromotor. The micromotor is composed of two concentric glass cylinders. The gap between the cylinders is filled with liquid crystal, 4-n-4\u2032-pentylcyanobiphenyl. A pulsed voltage from a function generator is applied between the cylinders. The homeotropic and planar orientation layers are treated on the facing surfaces of the outer and inner cylinders, producing a hybrid orientation profile between the cylinders.", + "texts": [ + " We extended the otion further to develop a driving motor of an actuator, aiming o broaden the application of liquid crystal actuators, such as the riving source of MEMS or lab-on-a-chip devices. In this work, we abricated a number of micromotors driven by the backflow of liqid crystals and investigated the driving performance of each in egard to parameters such as the dimensions of the micromotors nd the voltage and frequency of the applied electric field. . Experimental method The liquid crystal micromotor is composed of two concentric lass cylinders of different diameters (Fig. 2). As each cylinder otates relative to the other, the inner cylinder acts as a rotaion axis if the outer cylinder is fixed. The gap between the two ylinders is filled with a liquid crystal, 4-n-4\u2032-pentylcyanobiphenyl 5CB), which exhibits a nematic state in the temperature range 5 \u25e6C\u201335 \u25e6C [17]. The surfaces of the cylinders are coated with an ndium-Tin-Oxide (ITO) layer to form an electrode pair. On the ITO ayers, a horizontal and circumferential orientation layer covers the t o 3 o q t A d i w r s outer surface of the inner cylinder (if the orientation layer is horizontal and axial, the inner cylinder moves axially)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002837_b978-0-12-816721-2.00012-9-Figure12.7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002837_b978-0-12-816721-2.00012-9-Figure12.7-1.png", + "caption": "Fig. 12.7 Application of supporting structures for DED application [5]. This emerging DFAM capability provides a commercially relevant research opportunity.", + "texts": [ + " This production approach requires that additional DED material be deposited to accommodate a machining allowance to remove the surface defects inherent to DED systems. Furthermore, when compared with homogeneous billet, the anisotropy of DED materials must also be considered. This anisotropy is associated with the rapid and directional solidification inherent to metal fusion AM systems, and is discussed in the context of PBF in Chapter 11. The development of DED support structures remains a commercially relevant but underutilized DFAM research opportunity; however, the technical feasibility of such structures has been demonstrated (Fig. 12.7). These structures provide an opportunity for cost-reduction in commercial DED production, as well as providing innovative opportunities in design and certification. For example, frangible DED support structures provide an opportunity for the limitation of internal residual forces within manufactured structures, thereby providing a mechanism for certification and process determination, as well as limiting the (potentially damaging) stresses applied to the platen structure. Furthermore, unlike powder-bed systems, DED systems do not provide inherrent support of overhanging structures, and therefore the lack of robust supporting structures restricts the achievable manufacturability of the DED process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000954_0954406213516305-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000954_0954406213516305-Figure7-1.png", + "caption": "Figure 7. The framework of the customized lifting multiwavelet packet transform. Figure 9. The bearing test bench.", + "texts": [ + " However, aforementioned methods are not adequate for non-stationary signals and multi-component model. Iterative application of the Hilbert transform was adopted to demodulate multi-dimensional output signal in this work. The approach was proved with the sensational accuracy for stationary and transient signal and multi-component extraction, and the procedure of algorithm was proposed in the paper by Gianfelici et al.32,33 The procedure of customized-lifting multi-wavelet packet transform A framework of the customized-lifting multi-wavelet packet transform is given in Figure 7, which includes five steps for the overall procedure. Step one is to acquire the vibration signals by experiment. Step two is the pre-processing input signals by balancing method in order to obtain the multi-dimensional signal. The improved adaptive redundant-lifting scheme-merged swarm fish algorithm is applied to the multi-dimensional signal in step three. Step four is the post-processing of output signals by iterative application of the Hilbert transform for the sake of acquiring multiple fault features in various output channels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000677_aeit.2014.7002051-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000677_aeit.2014.7002051-Figure6-1.png", + "caption": "Fig. 6: Schematic representation of the Hely prototype used as a screwing machine.", + "texts": [ + " The blocks highlighted in blue allow the control of the linear motion, while the blocks highlighted in red are used for the control of the rotational movement of the Hely secondary. In addition, with the blocks included in the area highlighted in green it is possible to enter several parameters required for the specific application, i.e. spin, load torque, turn-on and turn-off times for each electric drive, length and degrees of displacement for the related applied load. In the first simulation, the Hely prototype is used as a screwing machine, as shown in fig. 6. With the adoption of the control panel, it is possible to control with high precision both the depth of the screw and the spin speed, as required for the proposed application. More in particular, after setting the values of the load torque and the load force, both linear and rotational motions are applied to the load. As an example, in the first 10 seconds of the simulation a linear movement is applied by driving LI with a reference speed of 0.001 m/s, while from 15 seconds to 45 seconds a rotary movement is imposed by driving RI with a reference speed of 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.40-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.40-1.png", + "caption": "Figure 9.40 A humanoid manipulator in its front and auxiliary views.", + "texts": [ + "\u0307?2 + \ud835\udf0e\u20323d4?\u0307?3 = \u2212(v\u22171s\ud835\udf032 + v\u22173c\ud835\udf032) (9.625) Equation (9.625) expresses the motion freedom induced in the joint space. Thus, as long as the task-space compatibility condition is satisfied, one of ?\u0307?2 and ?\u0307?3 can be assigned a finite arbitrary value and the other one can be found from Eq. (9.625). In this case, the manipulator can have only one kind of motion singularity, which is the same as the second kind of motion singularity explained in Section 9.7.8.1. A humanoid manipulator is shown in Figure 9.40 with an 8-DoF kinematic model including the rotation of the torso about the vertical axis. As a spatial manipulator, it has two degrees of redundancy. The significant points of the manipulator are as follows: O: Torso center point (origin of the base frame) S: Shoulder point, E: Elbow point, R: Wrist point, P: Tip point (a) Joint Variables \ud835\udf031, \ud835\udf032, \ud835\udf033, \ud835\udf034, \ud835\udf035, \ud835\udf036, \ud835\udf037, \ud835\udf038 They are the rotation angles about the joint axes. They are all shown in Figure 9.40. (b) Twist Angles \ud835\udefd1 = 0, \ud835\udefd2 = \u2212\ud835\udf0b\u22152, \ud835\udefd3 = \ud835\udf0b\u22152, \ud835\udefd4 = \u2212\ud835\udf0b\u22152, \ud835\udefd5 = \ud835\udf0b\u22152, \ud835\udefd6 = \u2212\ud835\udf0b\u22152, \ud835\udefd7 = \ud835\udf0b\u22152, \ud835\udefd8 = \u2212\ud835\udf0b\u22152 (c) Offsets d1 = 0, d2 = OS, d3 = 0, d4 = SE, d5 = 0, d6 = ER, d7 = 0, d8 = RP The nonzero offsets are given the following special names. d2: Shoulder length (shoulder offset), d4: Upper arm length, d6: Front arm length, d8: Hand length (tip point offset) (d) Effective Link Lengths b0 = 0, b1 = 0, b2 = 0, b3 = 0, b4 = 0, b5 = 0, b6 = 0, b7 = 0 As noticed, they are all zero. The actual lengths of the members are represented by the offsets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001824_1.4033034-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001824_1.4033034-Figure1-1.png", + "caption": "Fig. 1. Parameters of the the system in the central position of the wheelset. The parameters of the local geometry at the nominal contact point (h, Rw, Rr and the curvature derivatives) can be calculated from the profile curves cw and cr.", + "texts": [ + " In Section 4, formulae for the nonlinearity factor are derived for different types of wheel and rail profiles. In Section 5, the relationships are derived between the nonlinearity factor and some related parameters of the literature. In Section 6, the accuracy of the approach of rolling radius difference is also tested by means of the nonlinearity factor of the kinematic oscillation. In this section, we present the model of a single wheelset rolling on a straight track. The parameters of the system are given in the central position of the wheelset (see Fig. 1). The nominal Ac ep te d Man us cr ip t N ot C op ye di te d Downloaded From: http://computationalnonlinear.asmedigitalcollection.asme.org/ on 03/21/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Rw,Rr Rw,Rr R\u2032w,R \u2032 r R\u2032\u2032w,R \u2032\u2032 r Fig. 2. Three classes of profiles. Left panel: conical wheel with square rail. Middle panel: profiles with constant curvatures. Right panel: profiles with varying curvatures. In case of curved profiles, the necessary parameters of the local geometry are listed (see Tab. 1). rolling radius is denoted by r, and the lateral half-distance between the contact points is denoted by b. The wheel and the rail profiles around the nominal contact point are described by the functions cw(uw) and cr(ur), respectively (see Fig. 1). The profiles are defined so that uw = 0 and ur = 0 correspond to the nominal contact point, and cw(0) = cr(0) = 0. The nominal contact angle is denoted by \u03b4, that is, c\u2032w(0) = c\u2032r(0) = tan\u03b4 =: h, (1) where h is the inclination of the profiles at the central position, which is also called nominal conicity. Along the paper, the prime \u2032 denotes derivatives with respect to the axial coordinates uw and ur of the profiles. We assume that the profile curves cw and cr are at least four times differentiable at the nominal contact point, which is a reasonable assumption for worn profiles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003419_aim43001.2020.9158833-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003419_aim43001.2020.9158833-Figure3-1.png", + "caption": "Fig. 3: Top view of differential drive robot.", + "texts": [ + " , c7 are c1 = 2 ( Iwd + Imd ) + Ixx + ( Izz \u2212 Ixx \u2212mcl 2 ) cos2(\u03b2) +mcl 2, c2 = mwr 2 + Iwa + Ima\u03b3 2, c3 = mcl cos(\u03b1) sin(\u03b2), c4 = mcl sin(\u03b1) sin(\u03b2), c5 = mcl cos(\u03b1) cos(\u03b2), c6 = mcl sin(\u03b1) cos(\u03b2), c7 = ( Ixx +mcl 2 \u2212 Izz ) sin(2\u03b2). It is well known that Eq. (11) is only valid for a system without non-holonomic constraints. Therefore, before any standard state-space based control design can be applied, an alternative approach is needed to represent the motion and constraints of the system [16]. Under rolling and no-slip conditions that engender nonholonomic constraints, the kinematic equations of the differential drive robot shown in Fig. 3 are [17] I x\u0307r sin(\u03b1)\u2212 I y\u0307r cos(\u03b1) = 0 (12) I x\u0307r cos(\u03b1)\u2212 I y\u0307r sin(\u03b1) = r 2 ( \u03b8\u0307r + \u03b8\u0307l ) (13) \u03b1\u0307 = r 2b ( \u03b8\u0307r \u2212 \u03b8\u0307l ) (14) The above three equations can be written in matrix form as A(q)q\u0307 = 0, where A(q) = sin(\u03b1) \u2212 cos(\u03b1) 0 0 0 0 cos(\u03b1) sin(\u03b1) b 0 \u2212r 0 cos(\u03b1) sin(\u03b1) \u2212b 0 0 \u2212r . (15) Once we have equations for the constraints of the system, Eq. (11) is modified to account for k kinematic constraints as M(q)q\u0308 + V q\u0307 +H(q, q\u0307) +G = E\u03c4 +A(q)T\u03bb, (16) where \u03bb \u2208 Rk is a vector of Lagrange multipliers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003676_ecce44975.2020.9235378-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003676_ecce44975.2020.9235378-Figure3-1.png", + "caption": "Fig. 3. Discretization of a loop.", + "texts": [ + " In contrast with radial flux rotary machines with a PCB having an optimized shape [6], Nl independant PCBs separated by an insulation layer, are stacked one on top of each other. Each PCB is made of a double-sided copper clad with copper tracks printed on both side according to the layout presented in Fig. 2. A PCB, whose winding pitch equals to the magnet pitch of the PM \u03c4 , is characterized by Np, the number of poles and by Nt, the number of loops forming a pole. The shape of these loops is made of a succession of linear segments whose end positions are also parameters of the winding. As shown in Fig. 3, each segment is defined by the coordinates of its extremities. Two coordinate systems are used. The first is fixed to the stator, and the second is fixed to the mover. These are related thanks to the following coordinate transformation: (x\u2032, y\u2032, z\u2032) = (x\u2212 q, y, z) with q the relative position between the stator and the mover. Evaluating the performances of the actuator requires a modeling of its characteristics. In this section, an electromagnetic model of the actuator is presented. This latter allows to predict the magnetic properties of the actuator, namely the distribution of the magnetic flux density of the PM and the magnetic flux linked by the windings, this latter together with the current that flows in the windings being responsible of the force that the actuator can develop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000826_s1052618814050136-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000826_s1052618814050136-Figure5-1.png", + "caption": "Fig. 5.", + "texts": [ + " 4, IIIa) and we shall consider the parallel crank mechanism as the sliding kinematic pair in the other two kinematic chains containing rotary motors. Thus, altogether we get 16 ele ments. The number of 5 one degree of freedom kinematic pairs p5 is 17. W = 6 \u00d7 (16 \u2013 1) \u2013 5 \u00d7 17 = 5. This result is true: the number of degrees of freedom is five. Let us consider a mechanism with six degrees of freedom and we use formula (1) for spatial mecha nisms to determine the number of degrees of freedom of the manipulating mechanism (Fig. 5). Let us calculate the number of elements n in the mechanism and number of the one degree of free dom kinematic pairs p5. The spatial mechanism (Fig. 5) includes base 1, output element 2, working body 3 and three kinematic chains. Each kinematic chain contains a sliding motor 4, 4 ', 4 ''; initial rotary kine matic pair 5, 5 ', 5 ''; initial element of the parallel crank mechanism 6, 6 ', 6 ''; rotary kinematic pairs of the parallel crank mechanism 7, 7 ', 7 ''; terminal rotary kinematic pair 8, 8 ', 8 ''; terminal element of the parallel crank mechanism 9, 9 ', 9 '', intermediate elements of the parallelogram and terminal element of the kinematic pair 10, 10 ', 10 ''" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002359_j.msea.2016.08.078-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002359_j.msea.2016.08.078-Figure1-1.png", + "caption": "Fig. 1. Size and shape schematic of Micro-CT scanning specimen.", + "texts": [ + " The image processing and analysis software, CTAn V1.13 (Bruker microCT, Kontich, Belgium) was employed to perform 2D quantitative analysis of microstructure of TC6 titanium alloy. Furthermore, these experimental results serve as a reference to verify the accuracy of the following 3D microstructure model of TC6 titanium alloy obtained by laboratory X-ray microtomography. The specimen of TC6 titanium alloy for X-ray tomogram was machined by wire electrical discharge machining. The shape and size are shown in Fig. 1, the top portion is the cylindrical scanning area with \u00d80.5 1 mm, and the bottom portion is a cylinder with \u00d84 4 mm performing as a supportive platform. It should be noted that the axial direction of the cylinder specimen is paralleled to the rolling direction in this study. The dual-energy Micro-CT acquisition was performed by scanning the specimen twice with different X-ray filters and applied tube voltages. The specimen should not be moved between scans, whereas the image magnification and resolution should be the same [24]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001041_2013047-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001041_2013047-Figure2-1.png", + "caption": "Fig. 2. Sagittal representation of the 3D-EIP model.", + "texts": [ + " 1) in the inverted pendulum model to approach the CoM lateral displacements as: \u03b2s = sin\u22121 [ yOP\u0304 \u2212 y\u0308OC \u00b7 cos(\u03b1s) g \u00b7 s ] (4) The 3D approached CoM position in 0 was therefore defined by:[ OC\u2032 1 ] 0 = T 0,P\u0304 [ P\u0304C\u2032 1 ] P\u0304 (5) where T 0,P\u0304 = \u23a1 \u23a2\u23a2\u23a2\u23a3 cos\u03b1s sin \u03b1s sin\u03b2s sin\u03b1s cos\u03b2s xP\u0304 0 cos\u03b2s \u2212 sin\u03b2s yP\u0304 \u2212 sin\u03b1s cos\u03b1s sin \u03b2s cos\u03b1s cos\u03b2s 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a6 (6) The predicted lateral CoM trajectory was therefore influenced by the roll rod angle and the lateral CoP position dragged from experimental data. Besides, using the homogeneous matrix formalism, the pitch and roll angles of the inverted pendulum rod were introduced in a common mathematical model and the 3D CoM trajectory was predicted avoiding separate plane modeling. We considered that the CoM displacement during the period of double support described a sinusoid (Zijlstra & Hof, 1997) and was therefore assimilated to a single pendulum trajectory (Fig. 2). Let d be the rod length of the single pendulum and E\u0304 its mean double support pivot point (DPP) above the CoM (Fig. 2). At the transition instants between simple and double supports (i.e. IC ), the positions of the mean CoP, CoM and mean DPP are aligned (Fig. 2), and d was computed in proportion to the single and double support repartition as d = Sd Ss s, where Sd was the double support length and Ss the single support length. The approached CoM trajectory during successive single and double supports was computed: [ OC\u2032 1 ] 0 = \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 T 0,P\u0304 [ P\u0304C\u2032 1 ] P\u0304 \u2200t \u2208 [TO , IC ] T 0,E\u0304 [ E\u0304C\u2032 1 ] E\u0304 \u2200t \u2208 [IC , CTO ] (7) where E\u0304 = (E\u0304, xE\u0304, yE\u0304, zE\u0304) defined the local frame linked to the single pendulum. CTO denoted the instant of contralateral toe-off and T 0,E\u0304 the homogeneous matrix from frame 0 to frame E\u0304 as: T 0,E\u0304 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 cos\u03b1d sin \u03b1d sin \u03b2d sin \u03b1d cos\u03b2d xE\u0304 0 cos\u03b2d \u2212 sin\u03b2d yE\u0304 \u2212 sin \u03b1d cos\u03b1d sin \u03b2d cos\u03b1d cos\u03b2d 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6(8) and EC\u2032 = \u2212 dzE\u0304 (9) \u03b1d(t) and \u03b2d(t) were respectively the pitch and roll angles of the single pendulum rod during double support", + " The approached CoM trajectory was computed using the 3D Extended Inverted Pendulum (3D-EIP) model as: [ OC\u2032 1 ] 0 = \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 T 0,P\u0304 [ P\u0304C\u2032 1 ] P\u0304 \u2200t \u2208 [TO , IC ] T 0,E\u0304 [ E\u0304C\u2032 1 ] E\u0304 \u2200t \u2208 [IC , CTO ] (12) and E\u0304C\u2032 = \u2212\u0303dzE\u0304 (13) where the constant rod length of inverted pendulum (\u0303s) and single pendulum (\u0303d) were respectively defined during single and double supports as: 1. \u0303s was the distance between P (t = TO) and C\u2032(t = TO); 2. at t = IC , P , C\u2032 and E were aligned and \u0303d = Sd Ss \u0303s. The trajectory of the DPP during double support was deduced to the displacement of the CoP displacement and of a homotethie of center C\u2032 described by a Sd Ss ratio at CIC (Fig. 2) as \u03b4(OE) = Sd Ss \u03b4(OP). The 3D-EIP model was therefore based on a few spatiotemporal data: the rod lengths, the length of support phases, the CoP displacement, the pitch and roll rotation angles. These parameters were dragged from experimental data and introduced into the mathematical model previously defined. A Root Mean Square difference (RMSd) of the CoM trajectory calculation was performed between the reference multibody model and the 3D-EIP model with CoP/DPP translations or the model considering fixed CoP/DPP" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001119_20140824-6-za-1003.01266-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001119_20140824-6-za-1003.01266-Figure4-1.png", + "caption": "Fig. 4 Torque estimation equipment and experimental set-up", + "texts": [ + " It consists of two steps: the model estimation step and the torque estimation step. In the model estimation step, the model parameters are calculated from the sEMG, known joint torque and joint angle data. In the torque estimation step, the unknown joint torque is calculated from the sEMG and the joint angle data. In addition, step 2 performs under unknown external force condition. 4. EXPERIMENTS In order to obtain different levels of contraction and different joint torques a load is applied at the forearm with a winder and pulley device as can be seen on Fig. 4. Most of the load is due to the gravity force of load, the inertia of the arm and the mass of the forearm can be neglected compare to the load mass. The load is tied to the winder through pulley with steel wires. A board is fixed on the winder. The subject extends his/her forearm under the chosen constraint applied by the load. Two types of weights enable to obtain two different levels of contractions. The load is applied to the forearm, therefore the palmar flexion and the dorsal flexion affects slightly the sEMG measurement, in particular for the Brachioradialis. The subject\u2019s upper arm is fixed on the board to avoid the gravity of self-maintaining the arm to work around the elbow joint. Consequently the muscles only actively contribute to the elbow flexion or extension. With the joint angle \u03b8 [rad], the load torque \u03c4 [Nm] can be calculated as follows: \u03b8\u03b8\u03b8\u03c4 fIrgmr cos)( (5) r [m]:winder radius :winder moment of inertia : Weight mass The measuring device used for our experiment is detailed in Fig. 4. The sEMG are measured for the three surface muscles that actuate the elbow joint flexion/extension: Biceps, Triceps and Brachioradialis as shown in Fig. 5. sEMG from Brachialis is supposed to be identical as the one measured at the Biceps brachii. Since the relation between the parameters is linear is not taken into account explicitly, but through the Biceps brachii sEMG signal. A potentiometer is attached to the winder and outputs the joint angle. The surface electromyogram measurement system used is a \u201cMyon 320\u201d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.22-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.22-1.png", + "caption": "Figure 10.22 The PSFK of the manipulator.", + "texts": [ + " On the other hand, as indicated before, when this PSFK occurs, the angles \ud835\udf19\u2032 1 and \ud835\udf03\u2217k (for k = 1, 2, 3) become indefinite according to Eqs. (10.508), (10.518), and (10.528). Consequently, as also implied by the appearance of Figures 10.17 and 10.18 in accordance with Eqs. (10.537) and (10.538), all the six links of the lower legs become parallel to each other and therefore they can swing out of control with arbitrary values of the passive joint variables. Because of this uncontrollability, the mentioned PSFK must of course be avoided. This PSFK of the manipulator is illustrated in Figure 10.22, in which only the first and third legs of the manipulator are shown (again as if \ud835\udefd3 = \u2212\ud835\udf0b) for the sake of neatness. As verified in the previous position-domain analysis, the moving platform of the delta robot always remains parallel to the fixed platform. Therefore, C\u0302 = C\u0302(0,7) = I\u0302 and \ud835\udf14 = \ud835\udf14 (0) 7\u22150 = 0. Consequently, Eq. (10.473) for the location of the tip point (i.e. p) and Eqs. (10.475) and (10.476) for the major independent kinematic loops IKL-1 and IKL-2 happen to be sufficient to describe the kinematic behavior of the manipulator of this robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002902_j.cirpj.2020.01.001-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002902_j.cirpj.2020.01.001-Figure8-1.png", + "caption": "Fig. 8. Scheme of the channel for molten pool temperature regulation.", + "texts": [ + " The scheme of the experiment with the pyrometer and the equipment photo: (1) deposited layer; (2) wire feed mechanism; (3) electron beam oscillating area; (4) Raytek Marathon MR1SACF pyrometer; (5) vapor shield (pipe); (6) measurement area; (7) ELA-40I electron beam gun. Mathematical model with temperature feedback analysis Owing to the high molten pool temperature during the deposition process, it is necessary to use non-contact temperature measurement methods, such as pyrometer. The scheme for temperature control with feedback is shown in Fig. 8. Please cite this article in press as: D.A. Gaponova, et al., Effect of reheatin wire deposition method, NULL (2020), https://doi.org/10.1016/j.cirpj.20 A virtual pyrometer is located behind the center of the electron beam. The pyrometer signal, Tmes, is the feedback signal and is compared with the set temperature, Tset. The difference of these signals is fed into the controller of the beam current regulator, which acts on the bias supply source connected to the electron gun control electrode. The proposed method was developed into a mathematical model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003181_tase.2020.2993277-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003181_tase.2020.2993277-Figure2-1.png", + "caption": "Fig. 2. Mechanical structure of nonlinear-stiffness-compliant components.", + "texts": [ + " The pulley is fixed to the shaft of the gearbox, and the outer cylinder is connected with the pulley by wires, which eliminates backlash. The roller is fixed to the outer cylinder with a flange shaft. Three elastic components are installed evenly on the surface of the inner cylinder with fastening screws. In order to achieve the bidirectional actuation, the elastic component consists of two symmetrical elastic elements, with one end connected and the other end disconnected via a gap, as shown in Fig. 2. In equilibrium position, a roller is placed at the symmetry axis of the component, and the profiles of two elements are tangent with the roller, so there is no interaction force between the roller and the elastic components. The elastic components are deformed when there is an interaction force resulting in relative rotation between the inner cylinder and the outer cylinder. The relative rotational angle, namely, the deformation of the elastic components, reflects the external loads on the joint. The relative motion can be measured by the magnetic linear encoder, whose tap is wrapped to the surface of the inner cylinder due to its flexibility. The core structure of the LDNSA is the compliant mechanism, which determines the mechanical intrinsic compliance and nonlinear stiffness of the system. According to the law of \u201csmall load, low stiffness; large load, high stiffness,\u201d elastic components are designed to have the characteristic of nonlinear stiffness. As shown in Fig. 2, the main structure of a compliant mechanism consists of a roller and an elastic component that consists of two symmetrical elastic elements. The elastic element is composed of an elastic part and a contact part. The elastic part is a cantilever beam structure that generates deflection when an elastic restoring force is exerted on it. Authorized licensed use limited to: University of New South Wales. Downloaded on July 26,2020 at 22:20:41 UTC from IEEE Xplore. Restrictions apply. The elastic part generates a deflection angle when the roller stresses the contact part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001129_iraniancee.2013.6599611-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001129_iraniancee.2013.6599611-Figure2-1.png", + "caption": "Fig. 2: Schematic diagram of the linear two- mass system with backlash", + "texts": [], + "surrounding_texts": [ + "In this section a two-stage identification algorithm is proposed to identify all unknown parameters of the system sequentially. Motor physical parameters ( , ), load physical parameters ( , ), load torque disturbance ( ) and physical parameters of the shaft ( , ) are the unknown parameters of the system that will be identified by the proposed two-stage identification algorithm. Deriving proper linear in parameter models for indirect adaptive position control and identifying all unknown parameters of the backlash system is the main goal of this section. In the following, the two-stage identification scheme is described." + ] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure4-1.png", + "caption": "Fig. 4. The reduced models for the wave and spiral channels.", + "texts": [ + " This model is used for the modelling of the machine with squiggly ducts in the end region, but also for the wave channel as a comparison to the 2-way coupled model. In this way we could compare the result from the different program packages. See the system model in Fig. 3. The reduced thermal models (in ANSYS Fluent) consist of the cooling system with the casing and the cooling duct with water, and a heat power boundary condition representing motor loss on the inside of the casing. The models are used to yield convective heat transfer coefficients at different flow rates and give input to the LPN-model. See models for the wave and spiral duct cases in Fig. 4. 868 Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 15,2021 at 16:48:44 UTC from IEEE Xplore. Restrictions apply. Outer boundary conditions of the thermal models are set on the inlet and outlet of the cooling duct and on the outer boundaries of the casing, cooling duct and shaft. The inlet flow rate and temperature are chosen as in [3] for the reason of comparison. The effect of different flow rates, and choice of boundary conditions are examined with the various models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001720_el.2015.1370-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001720_el.2015.1370-Figure2-1.png", + "caption": "Fig. 2 Deployment point compensations a Compensation in one cluster using dumax b Deployment on maximum range dumax to reduce overall duration", + "texts": [ + " Therefore (5) can be expanded and approximated as follows: While the deployment point is calculated by (6), the maximum travelling distance contraint of a UAV is not considered yet. Therefore all displacements from the deployment point to the corresponding tasks should be verified. To reduce computation time, if any radius \u03d5s is longer than No. 21 pp. 1650\u20131652 dumax, then the corresponding cluster is divided into two clusters iteratively until all divided clusters satisfy the condition. Let ms max be the farthest task from ws, and c be the angle from the moving direction of the carrier to ms max as shown in Fig. 2a. If \u2016wsms max\u2016 . dumax, then the position of ws should be shifted by \u03b4d as follows: dd = \u2016wsm s max\u2016 cosc\u2212 (dumax) 2 \u2212 (\u2016wsms max\u2016 sinc)2 \u221a (7) On the other hand, assume there are three clusters as shown in Fig. 2b. Then the carrier does not have to reach the next deployment point ws+1 at the same time as the previous deployed UAVs reach their task locations unless the next deployment point is the last one. Therefore the carrier should instead deploy UAVs at its maximum range dumax as the UAVs would have reached the task locations anyway by the time the carrier arrived at ws+2. The overall procedures are designed as follows. Algorithm 1: Dynamic Programming 1: Initialise Cont true, IdxDIV 1, IdxMAX 1, Elapsed T TMAX, PrevDeploymentPos (xRC , yRC ) 2: while Cont is true do 3: DIVIDE_TASK(PrevDeploymentPos, Task, IdxDIV) 4: IdxMAX\u2190 IdxMAX + 1 5: for i\u2190 1 to IdxMAX do 6: (centre[i], rad[i])\u2190 GET_MINCIRCLE(i) 7: if rad[i] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002198_978-3-319-14194-7_4-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002198_978-3-319-14194-7_4-Figure3-1.png", + "caption": "Fig. 3 Type-2 Gaussian fuzzy sets with uncertain standard deviation", + "texts": [ + " In this layer each input variable xi from the I-dimensional input space of the network is quantized into n discrete sub-regions according to a defined resolution. A possible quantization of the input (sensor) space for two dimensional CMAC network is presented in Fig. 1. Layer 2. Fuzzification layer: The quantized in layer 1 input space is fuzzified here by predefined fuzzy sets with corresponding membership functions. A type-2 fuzzification of the input space of the fuzzy CMAC network is implemented in this investigation. Interval type-2 fuzzy sets (IT2-FSs), presented by type-2 Gaussian membership functions with uncertain standard deviation (Fig. 3), have been used into the second layer of the P-T2FCMAC architecture. The lower and upper membership functions for type-2 Gaussian membership functions with uncertain deviation are expressed by: lij\u00f0xi\u00de \u00bc exp 1 2 \u00f0xi cij\u00de2 r2ij ! \u00f011\u00de l ij \u00f0xi\u00de \u00bc exp 1 2 \u00f0xi cij\u00de2 r2ij ! \u00f012\u00de where cij \u00f0i \u00bc 1 . . . I; j \u00bc 1 . . . J\u00de is the mean value of the jth fuzzy set for the ith input signal; rij and rij are deviations of the upper and lower membership functions of the jth fuzzy set for the ith input signal", + " Neurons in this layer are represented by the firing strength of the corresponding rule Rr calculated as the T-norm product of the degrees of fulfillment for both membership functions (upper and lower) respectively: wr \u00bc l~A1\u00f0x1\u00de l~A2\u00f0x2\u00de l~AI\u00f0xI\u00de wr \u00bc l~A1 \u00f0x1\u00de l~A2 \u00f0x2\u00de l~AI \u00f0xI\u00de \u00f013\u00de Moreover, memory element selection (association) vectors aTr ought to be determined in this layer, according the idea proposed in [6]. It was mentioned before that the activated cells formed by shifting units are called receptive fields or hyper-cubes [26]. Each location corresponds to a fuzzy association given by (10). For the P-T2FCMAC network with two inputs, introduced in Fig. 2, type-2 fuzzy sets with three Gaussian membership functions (Fig. 3) are imposed on each input variable xi. Accordingly, there are 9 fuzzy rules with 9 association vectors, aTr , (r = 1, 2,\u2026 9). The reduction of the association vectors\u2019 number is obvious here. In the CMAC network there are 16 vectors, while fuzzy reasoning determines 9 association vectors. The logical operation \u2018OR\u2019 is performed on all possible (in the same region) association vectors in CMAC. Layer 4. Fuzzy post-association layer: It is also called weight memory space. The layer performs normalization of the firing strengths calculated in the previous layer [26]: ~wr \u00bc wrPN r\u00bc1 wr ; ewr \u00bc wrPN r\u00bc1 wr \u00f014\u00de Furthermore, the output weights of this layer correspond to the weighting column vector P of the P-T2FCMAC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002038_ijeoe.2016040101-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002038_ijeoe.2016040101-Figure1-1.png", + "caption": "Figure 1. Earth fixed and Body fixed frames", + "texts": [ + " in translational and rotational along x, y, & z axes. The notations used for modeling are in SNAME nomenclature (see Table 1). The general motion of vehicle is described by following vectors. \u03b7 \u03b7 \u03b7 \u03b7 \u03b7 \u03d5 \u03b8 \u03c8= = =[ , ] , [ , , ] , [ , , ] ;1 2 1 2 T T T T Tx y z (1) v v v v u v w v p q rT T T T T= = =[ , ] , [ , , ] , [ , , ] ; 1 2 1 2 (2) \u03b7 denotes position and orientation in earth fixed frame, v denotes linear and angular velocities in body fixed frame, \u03c4 denotes forces and moments acting on vehicle in body fixed frame (Figure 1). In two coordinate system the equations of motion of AUV can be written as in equation (4) and (5). \u03b7 \u03b7= J v( ) (4) Mv C v v D v v g + + + =( ) ( ) ( )\u03b7 \u03c4 (5) Here v u v w p q r T= [ , , , , , ] is a vector of linear and angular velocities, \u03c4 = [ , , , , , ]X Y Z K M N T represents forces and moments, M is inertial matrix, C(v) is Coriolis and centripetal matrix, g(\u03b7) is gravitational matrix, D(v) is damping matrix. J(\u03b7) in equation (4) is transformation matrix using Euler Angle. Nonlinear equations of motion in 6 DOF giving forces and moments are obtained from equation (5) and are as follows (Issakhov, 2014): m u vr wq x p q y pq r z pr q X G G G \u2212 + \u2212 + + \u2212 + + =( ) ( ) ( )2 2 (6) m v wp ur y r p z qr p x qp r YG G G \u2212 + \u2212 + + \u2212 + + =( ) ( ) ( )2 2 (7) m w uq vp z p q x rp q y rq p Z G G G \u2212 + \u2212 + + \u2212 + + =( ) ( ) ( )2 2 (8) I p I I qr r pq I r q I pr q I m y w u xx zz yy xz yz xy G + \u2212( ) \u2212 +( ) + \u2212( ) + \u2212( ) + \u2212 2 2 q vp z v wp ur KG+( )\u2212 \u2212 +( ) = (9) I q I I rp p qr I p r I qp r I m z u yy xx zz xy zx yz G + \u2212 \u2212 + + \u2212 + \u2212 + \u2212 ( ) ( ) ( ) ( ) ( 2 2 vr wq x w uq vp M G + \u2212 \u2212 + =) ( ) (10) I p I I qr r pq I r q I pr q I m y w u xx zz yy xz yz xy G + \u2212( ) \u2212 +( ) + \u2212( ) + \u2212( ) + \u2212 2 2 q vp z v wp ur KG+( )\u2212 \u2212 +( ) = (11) First three equations i e. equations no. (6), (7), & (8) are for translational motion and remaining are for rotational motion. 3.2. AUV Subsystems Complete mathematical model of AUV is divided into three subsystems: 1. Speed system 2. Depth system 3. Steering system All the hydrodynamic coefficients and parameters used in depth and steering state space models are of REMUS-100 AUV (Figure 1). It is in Myring hull shape with forward velocity U=1.54m/s (knots), mass m =30kg. The parameter values are as in Table 2 (Prestero, 2001). Smooth operation of AUV is mainly dependent on the working of thrusters. In AUV there are two vertical and two horizontal thrusters. For attaining a specified depth, the thrusters of AUV play major role. During operation if one of the vertical thrusters fails partially or completely then AUV becomes unstable. Hence under failure of vertical thrusters model of depth subsystem is of importance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003088_tia.2020.2992579-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003088_tia.2020.2992579-Figure9-1.png", + "caption": "Fig. 9. Flux line in steady state. The steady state torque of DSLSPMV machine with", + "texts": [ + " It can be seen that the motor can produce large enough torque when stator winding pole-pair number is changed from 1 to 2. The average torque in steady state is 51.7Nm and the rotor-locked average torque is about 100Nm, the starting torque multiple is 1.93 for proposed 6-statorslot/5-PM-pole-pair LSPMV machine. III. CASCADING METHOD Cascading method is the other method to acquire selfstarting capability for PMV machine. The principle of cascading method can be seen as a cascade of an IM and a PMV machine. The IM-PMV machine system is shown in Fig.9 (a). The cascading machine system starting process can be described as follows: Firstly, rotor is speeded up as soon as the IM stator winding is connected to the power grid; then, the PMV machine winding is connected to the power grid to synchronize when the rotor speed is near the IM synchronous speed; finally, the IM stator winding is disconnected from the power grid after PMV machine synchronization. In order to reduce the volume, the DSLSPMV machine is proposed based on IM-PMV machine system. The DSLSPMV machine slide view is shown in Fig.9. (b). The squirrel cage of IM is embedded in the inner circumference of the rotor and the stator of IM is as the inner stator of DSLSPMV machine. The PMV machine and the IM share one rotor. The coupling of outer stator winding and the squirrel cage can be nearly eliminated when the rotor yoke length is large enough. Therefore, the starting performance and steady state performance can be decoupled, and the steady state performance is the same as the regular PMV machine. The wound winding also can be selected as the rotor starting unit when the operation braking toque is too heavy", + " If wound winding is adopted, the rotor wound winding pole-pair number Pr and inner stator winding pole-pair number PIM should satisfy: = r IMP P (10) If the synchronous speed of IM is much lower than synchronous speed of PMV machine, the rotor cannot be pulled into synchronization when the outer stator winding is connected to the power grid. In order to make sure the rotor speed can be accelerated to synchronous speed of VPM machine, the PIM and Pe should be satisfied as: e 60 60 IM f f P P (11) The circuit of DSLSPMV machine is shown in Fig.8 (a), and the dimension parameters are given in Table IV. It can be seen from Fig.9 (a) that the outer winding APMV, BPMV, and CPMV are connected with a time relay SPMV. The inner winding AIM, BIM, and CIM are connected with a time relay SIM. The whole starting and operation process can be divided into three parts including early starting process, synchronization period and steady state. Fig.9 (b) shows the early starting process. When the DSLSVPM machine begins to start, the switch SIM is off and SPMV is on, meanwhile, the inner stator winding is connected to the power grid and the rotor is accelerated. The winding connection of synchronization period is shown in Fig.9 (c). After rotor speed is near the synchronous speed, the SIM is on, and the rotor starts to synchronize. Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 03,2020 at 06:44:11 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Fig.9 (d) shows the winding connection of steady state start. The SIM is off after the motor is pulled into synchronization. Finally, the whole DSLSVPM machine operates as a regular PMV machine. The rotor yoke length determines the kind of rotor starting unit. Fig.9 shows the flux line of proposed DSLSPMV machine with rotor wound winding in steady state, as can be seen, the effect of decoupling is obvious. By choosing a proper rotor wound winding pole-pair number, the operation braking torque even can be eliminated. The squirrel cage is also can be chosen as rotor starting unit due to the weak coupling in steady state. different rotor starting unit is shown in Fig. 10. It can be seen that the torque waveforms nearly coincide because only a few flux lines go through the rotor starting unit as shown in Fig.9, and the average torque is 65.2Nm. Fig.11 shows the torque-speed curve of DSLSPMV machine when switch SPMV is off and SIM is on. The mechanical characteristic of pump load is also given. The pump load torque variation with speed can be expressed as: 2(1 )load ratedT T s (12) where Tload and Trated is pump load torque and rated torque respectively. It can be seen in Fig.11 that the electromagnetic torque curve intersects with load curve at point \u201cA\u201d. When the torque-speed curve of electromagnetic torque (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003406_asjc.2398-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003406_asjc.2398-Figure4-1.png", + "caption": "Figure 4 and the motion of the robot system is in a vertical plane. q1 and q2 are the angular displacement of links 1 and 2, respectively. m1 and m2 are the mass of links 1 and 2, respectively. \u03c41 and \u03c42 are the control torque in joints 1 and 2, respectively. l1 and l2 are the length of links 1 and 2, respectively. lc1 and lc2 are the", + "texts": [ + " Thus (32a-b) can be rewritten H0 11 +\u0394H11 \u20acq1 + H0 12 +\u0394H12 \u20acq2 = h _q2\u2212c1\u00f0 \u00de _q1 + h _q1 + _q2\u00f0 \u00de _q2\u2212g1 + \u03c41, \u00f033a\u00de H0 21 +\u0394H21 \u20acq1 + H0 22 +\u0394H22 \u20acq2 = \u2212h _q21\u2212c2 _q2\u2212g2 + \u03c42: \u00f033b\u00de One can find the matrix form as M _Q=N _Q+D+U, \u00f034\u00de where M=M0 +\u0394M, M0 \u00bc 1 0 0 0 0 H0 11 0 H0 12 0 0 1 0 0 H0 21 0 H0 22 26664 37775, \u0394M\u00bc 0 0 0 0 0 \u0394H11 0 \u0394H12 0 0 0 0 0 \u0394H21 0 \u0394H22 26664 37775,Q\u00bc Q1 Q2\u00bd T ,Q1 \u00bc q1 _q1\u00bd T,Q2 \u00bc q2 _q2\u00bd T, N\u00bcN0\u00fe\u0394N, N0 \u00bc 0 1 0 0 0 \u2212c1 0 0 0 0 0 1 0 0 0 \u2212c2 26664 37775,\u0394N\u00bc 0 0 0 0 0 h _q2 0 h _q1\u00fe _q2\u00f0 \u00de 0 0 0 0 0 \u2212h _q1 0 0 26664 37775,U\u00bc 0 u1 0 u2\u00bd T, u1 = \u03c41, u2 = \u03c42, and D\u00bc 0 \u2212g1 0 \u2212g2\u00bd T: Then, 34 can be rewritten as _Q=A0Q+G0U+\u0394, \u00f035\u00deFIGURE 4 Two-link manipulator robot system where A0 = M\u22121 0 N0, G0 =M\u22121 0 , and \u0394=M\u22121 0 D+\u0394NQ\u2212\u0394M _Q = \u03941 \u03942 \u03943 \u03944\u00bd T: Let a reference model of the two-link manipulator robot system be _Q =AmQ +GmUc, \u00f036\u00de where Q = Q 1 Q 2\u00bd T is the desired state vector, Q 1 = q 1 _q 1 T and Q 2 = q 2 _q 2 T , q 1, _q 1, q 2, and _q 2 are the control objectives, Am 2R4\u00d7 4, Gm 2R4\u00d7 4, and Uc = 0 u1c 0 u2c T , and u1c and u2c are the command inputs for joints 1 and 2, respectively. Let the control input U of the MRAC be U= \u03a6\u0302fUc + \u03a6\u0302bQ\u2212G\u22121 0 \u0394\u0302, \u00f037\u00de where \u03a6\u0302f = 0 0 0 0 0 \u03d5\u0302f1 0 0 0 0 0 0 0 0 0 \u03d5\u0302f2 26664 37775, \u03d5\u0302f1 and \u03d5\u0302f2 are the forward control gains, \u03a6\u0302b = 0 0 0 0 \u03d5\u0302b11 \u03d5\u0302b12 \u03d5\u0302b13 \u03d5\u0302b14 0 0 0 0 \u03d5\u0302b21 \u03d5\u0302b22 \u03d5\u0302b23 \u03d5\u0302b24 26664 37775, \u03d5\u0302b11 \u2026\u03d5\u0302b24 are the feedback control gains, and \u0394\u03022R4\u00d7 1 is the estimated disturbance vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001814_2016-01-1132-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001814_2016-01-1132-Figure2-1.png", + "caption": "Figure 2. Premium Tripot-type joint", + "texts": [], + "surrounding_texts": [ + "The powertrain of cars and light trucks has been evolving to reduce fuel consumption. Recent fuel efficiency improvements have been accomplished through various techniques including engine downsizing, turbocharging, stop-start systems and transmissions with more gears. Mechanical and friction losses in the engine still account for most of the mechanical energy losses in a vehicle, followed by losses in the transmission, aerodynamic drag and tire rolling resistance. Mechanical energy losses due to driveline and chassis components (i.e. differential, driveshafts, halfshafts, wheel bearings), in a typical compact car are in the order of 2% of the total mechanical energy losses [1]. Therefore, the contribution of the Halfshafts (HS) to fuel economy, with respect to other systems and sub-systems in the vehicle, is small. It is estimated that a 0.1 percentage point\nmechanical efficiency improvement in a set of Halfshafts would result in fuel economy improvements of approximately 0.1% or less, under typical highway conditions.\nConstant Velocity Joints (CVJ\u2019s) can be classified according to its architecture in two large groups: Ball-type joints (i.e. Rzeppa, Cross-Groove, Double-Offset) and Tripot-type joints (i.e. Standard and Premium Tripot). Relative motion, among the parts of a CVJ, has a rolling and a sliding component. Most of the frictional losses in CVJ\u2019s are caused by the sliding component. Therefore, if an improvement in mechanical efficiency is desired, two strategies (or the combination of both) can be followed which are:\n\u2022 Reduce the amount of sliding through kinematics, i.e. cage balancing through track-geometry [2];\n\u2022 Reduce the frictional forces through a tribological approach (grease and texture).\nIn this work, differences in mechanical efficiency among different types of CVJ\u2019s are quantified, as well as the improvement in mechanical efficiency that can be achieved through the right selection of grease for the application.", + "Joint type, grease, bending angle, torque and rotating speed influence the amount of mechanical losses in a CVJ. Thus, an extensive study was performed to understand the effect of the variables listed above. Nine halfshaft configurations were used, listed in Table 1, each\nhalfshaft configuration included three assemblies, minimum. The test specimens comprehended three CVJ global manufacturers, eight different styles of joints in production and four types of grease known by the authors and used in halfshafts supplied to OEM\u2019s, plus at least four other types of grease present in original aftermarket parts. The joints were tested individually and as halfshaft assemblies, to verify the individual efficiencies factored into the measured efficiency of the halfshaft. The halfshafts not built by the authors were obtained as aftermarket parts through licensed automotive dealers in the U.S. and Europe and were not modified before or during the conducted study.\nTest parameters were defined such that conditions representative of city driving, suburban driving and highway cruising could be captured, shown in Table 2.\nThe test conditions in Table 2 were performed at joint bending angles of 5\u00b0, 10\u00b0 and 15\u00b0.\nAll CVJ\u2019s go through a break-in period in which grease additives get activated on contact areas and asperities reduced due to components interaction. All parts in this study ran on a test bench, with torque applied, for a period of time, to ensure the joints were broken-in prior to measuring its efficiency.", + "There are several approaches to quantify mechanical efficiency (i.e. measurement of dissipated heat). Biernmann [3] lists and addresses the limitations of several known techniques. A limitation when quantifying mechanical losses through torque measurement at the input and output shafts of a mechanical system is the accuracy of the load (torque) cells used in the measurement and offset errors. To minimize these limitations the Central Limit Theorem was used, treating the measurements in steady state conditions as populations of data points rather than snapshots in time. The load cell sensitivities were verified for correlation at zero bending angle; establishing a zero reference for each test condition while rotating the halfshaft assembly, prior to applying the test torque and changing bending angle. This allowed minimization of the influence of the test stand in the mechanical efficiency measurements. Figure 3 shows a torque signal during a typical test event and the regions that were used to establish a zero reference and the deltas in input and output torques.\nMathematically, considering the speeds at the input and output shafts of a CVJ are the same, the approach can be described as follows:\nT_in0 = Mean baseline input torque when the part is rotating at zero bending angle, no test torque applied\nT_out0 = Mean baseline output torque when the part is rotating at zero bending angle, no test torque applied\nT_in1 = Mean test torque applied at input shaft\nT_out1 = Mean torque at output shaft\n(1)\n(2)\n(3)\nThe estimated accuracy of the measurement system used in this study for mechanical efficiency, based on the characteristics of the load cells employed, is +/- 0.03 percentage points, which is similar to the accuracy of the measurement system, based on heat losses, described by Biermann.\nThe data from the test conditions listed in Table 2 was used to generate efficiency maps of the different joints and halfshafts configurations. Figure 4 shows two efficiency maps, one for group A (top) and one for group B (bottom), at a bending angle of 5\u00b0. It can be observed the influence of torque and speed in mechanical efficiency, as well as the improvement that can be achieved through grease technology.\nThe hardware in groups A and B corresponds to the same design and was manufactured at the same time (single batch), the only difference within the former groups of parts is the grease used for lubrication." + ] + }, + { + "image_filename": "designv11_34_0000346_asjc.873-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000346_asjc.873-Figure2-1.png", + "caption": "Fig. 2. Schematic of the quadrotor.", + "texts": [ + " The remainder of this paper is organized as follows: in Section II, the quadrotor model description is presented; in Section III, the decoupled robust controller is designed; robust tracking properties are proven in Section IV; experimental results on the quadrotor rotorcraft are presented in Section V; and conclusions of this paper are drawn in Section VI. Let \u03b1 = {Ex, Ey, Ez} denote the earth-fixed inertial frame and \u03b2 = {Ebx, Eby, Ebz} the body-fixed frame with its origin in the center of the mass of the helicopter, as shown in Fig. 2. The changes of the lift forces produced by the four rotors lead to the movements of the quadrotor. In this paper, the quadrotor rotorcraft flies with an \u201cX\u201d configuration in order to increase the maneuverability and agility. As depicted in Fig. 3, the body fixed frame of the \u201cX\u201d configuration quadrotor is defined differently from that of the regular configuration. In regular configuration, the longitudinal or latitudinal motion can be achieved by changing the angular velocities of only two rotors", + " In the \u201cX\u201d configuration, however, all of rotational velocity of the four rotors must be changed in order to achieve the desired torque to move in the \u00a9 2014 Chinese Automatic Control Society and Wiley Publishing Asia Pty Ltd longitudinal or latitudinal direction. The difference of the counter-torque generated between the pair (the front left and back right rotors) and the pair (the front right and back left rotors) can result in the motion in the yaw direction. The change of total thrust of the four rotors causes the vertical motion of the quadrotor. Let \u03be = [\u03bex \u03bey \u03bez]T denote the position of the helicopter mass center in the inertial frame \u03b1 and v v v vx y z T= = [ ] \u03be the velocity of its center in Frame \u03b1. As depicted in Fig. 2, the orientation from Frame \u03b1 to Frame \u03b2 can be described by three Euler angles: the pitch angle \u03b8, the roll angle \u03d5, and the yaw angle \u03c8 [20]. This quadrotor has four control inputs ui(i = 1, 2, 3, 4) to change the body fixed frame torques about the axes Ebx, Eby, and Ebz, and the total lift force, respectively. The quadrotor system is considered to be four single-input single-output (SISO) subsystems with couplings among them: the longitudinal, latitudinal, vertical, and yaw subsystems. The control inputs and outputs for the four subsystems are u1 and vx (the longitudinal velocity), u2 and vy (the latitudinal velocity), u3 and \u03c8, and u4 and \u03bez (the height)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002135_amc.2016.7496392-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002135_amc.2016.7496392-Figure9-1.png", + "caption": "Fig. 9. Forces and reaction forces on holds", + "texts": [ + " In the case of grasping a bracket holds, the hand can pull the hold for not only vertical downward direction but also obliquely downward direction. Hence, the reaction force direction is not only vertical upward direction but also obliquely upward direction. On the other hand, at a crimp holds, only fingertip or toe can be put on, and the force direction is restricted. In the case of frictionless holds, hand or foot can push to the normal direction of the hold surface. Finally, for sloper holds, hand relys on only friction. The fan-shapes drawn in Fig.9 express the generable direction of forces, we call them hold-cones. It is similar to friction cone used in robotics field. The center angles of the hold-cone depends on the shape of holds. To keep equilibrium state of climbing robot, the sum of total forces and the sum of moments should be zero. Suppose that a robot has N limbs and put n(\u2264 N) limbs on some holds. Here, the word put means a generic term to represent grasping holds, hooking fingers on holds, pushing holds by palms, and riding holds on toes or heels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002831_ls.1497-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002831_ls.1497-Figure5-1.png", + "caption": "FIGURE 5 Deformation vector plot for \u03b5 = 0.33 and L/ D = 0.66 [Color figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " The low rotational speed of the shaft drags less oil vapor bubbles into the inlet area, and when the film is converging, the liquid oil is squeezed and pushes the oil vapor bubbles out of the journal bearing and less cushioning effect in the convergent portion; thus, it formed almost the same pressure build-up with the cavitation model. Also, study with cavitation model defined the negative pressure zones in the divergent region of the bearing, and negative pressure of 0.471 MPa was witnessed due to gases dissolved in the lubricant released because of oil pressure in that region was less than the saturation pressure and enhanced the cavitation effect. However, the flow analysis with cavitation is extremely important in case of the bearings operating with higher speeds. Figure 5 illustrates the displacement vector plot of bearing deformation under the influence of hydrodynamic forces for the eccentricity ratio of 0.33. The observed 3D vector plot is same as pressure build-up that occurs in the TABLE 1 Details of the bearing and oil properties Engine type Three-cylinders inline engine Engine cubic capacity 1300 cc Rating 36 kW@2250 rpm Type of fuel used Diesel Journal radius 24 mm Bearing length 32 mm Radial clearance 30 \u03bcm Rotational speed range 2250 RPM Lubricant viscosity 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure5.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure5.1-1.png", + "caption": "Figure 5.1 Two rigid bodies connected with a joint.", + "texts": [ + " Companion Website: www.wiley.com/go/ozgoren/spatialmechanicalsystems 82 Kinematics of General Spatial Mechanical Systems variables together with their first and second rates that correspond to the position and motion specifications. The solutions are accompanied by the discussions about the multiple poses and the poses of position and motion singularities. Consider a rigid body system that contains the rigid bodies o, a, b, ..., z as its members. Two of the members (a and b) of the system are shown in Figure 5.1. The members a and b are connected to each other with a joint denoted as ab or ba, which is also called a kinematic pair. The kinematic pair ab = ba consists of two mating kinematic elements, which are ab fixed on a and ba fixed on b. Usually, a member a of the system is also called a link by conceiving it as a bridge that links two or more joints. When it is called so, it is denoted as a. However, it is to be noted that, from a logical point of view, a rigid body may be called a link if it contains at least two joints", + "\u201d This usage is based on the fact that saying \u201cthe location of a\u201d is an abbreviated way of saying \u201cthe position of a representative point of a\u201d and saying \u201cthe orientation of a\u201d is equivalent to saying \u201cthe angular position of a.\u201d In a similar sense, the word \u201cpose\u201d is also used sometimes as an alternative to the word \u201cposition.\u201d However, they are actually not quite the same, because the word \u201cpose\u201d implies only the combination of the words \u201clocation\u201d and \u201corientation.\u201d Besides, the word \u201cpose\u201d has also a meaning close to that of the word \u201cposture.\u201d Kinematics of Rigid Body Systems 85 Referring to Figure 5.1, the relative orientation ofb with respect toa can be expressed by the following equation. C\u0302(a,b) = C\u0302(a,ab)C\u0302(ab,ba)C\u0302(ba,b) (5.1) In Eq. (5.1), C\u0302(a,ab) and C\u0302(ba,b) = C\u0302(b,ba)t are constant matrices. The matrices C\u0302(a,ab) and C\u0302(b,ba) describe the orientations of the kinematic elements ab and ba with respect to the body frames a and b, respectively. As for C\u0302(ab,ba), it happens to be a variable matrix, if the joint ab has rotational mobility. For such a joint, C\u0302(ab,ba) is a function of the joint variable or variables that describe the relative orientation of ba with respect to ab. Referring to Figure 5.1 again, the relative location of the origin Ob with respect to the other origin Oa can be expressed by the following equation. r\u20d7a,b = r\u20d7a,ab + r\u20d7ab,ba + r\u20d7ba,b (5.2) In Eq. (5.2), the relative position vectors are defined so that r\u20d7a,b = \u2212\u2212\u2212\u2192OaOb, r\u20d7a,ab = \u2212\u2212\u2212\u2192OaOab, etc. As seen in Figure 5.1, it happens that r\u20d7ba,b = \u2212r\u20d7b,ba. Equation (5.2), which is a vector equation, can also be written as a matrix equation in one of the relevant body frames a and b. Here, it is written in a as the following matrix equation. r(a)a,b = r(a)a,ab + r(a)ab,ba + r(a)ba,b \u21d2 r(a)a,b = r(a)a,ab + C\u0302(a,ab)r(ab) ab,ba + C\u0302(a,b)r(b)ba,b (5.3) Equation (5.3) can be written with more detail as follows: r(a)a,b = r(a)a,ab + C\u0302(a,ab)[r(ab) ab,ba + C\u0302(ab,ba)C\u0302(ba,b)r(b)ba,b] (5.4) In Eqs. (5.3) and (5.4), r(a)a,ab and r(b)ba,b = \u2212r(b)b,ba are constant column matrices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001123_amm.660.633-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001123_amm.660.633-Figure1-1.png", + "caption": "Fig. 1 Specification of Rollover Test Fig. 2 Specification of Quasi-Static Test", + "texts": [ + "1% of bus accident is caused by the factor of superstructure under rollover with highest percentage of fatalities occurrence (60%) [3]. The rollover test approach normally involve the lateral tilting test [4]. It is performed by locating a simplified vehicle on the tilting platform, with blocked suspension. The vehicle is lifted slowly to its unstable equilibrium position. This test can be further simplified by positioning the vehicle in unstable equilibrium at point of rollover. The contact position is set very close to the ground as depicted by Fig. 1. In quasi-static loading test, the load that applies to the beam around the body section is derived from the mass of the structural bays and elements that connecting them as shown in Fig. 2. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 128.210.126.199, Purdue University Libraries, West Lafayette, USA-01/06/15,21:04:51) The simulation process involved nine main stages as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002006_s11071-016-2788-z-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002006_s11071-016-2788-z-Figure3-1.png", + "caption": "Fig. 3 Link between swashplate and the cyclic pitch of Bell\u2013 Hiller bar", + "texts": [ + " Simulation and experimental results in hovering mode are displayed in Sect. 6. This paper is concluded in Sect. 7. 2 The mechanical structure of the main rotor As is shown in Fig. 1, the Thunder Tiger Raptor 50 is a levo-rotational helicopter with two hingeless blade. Figure2 shows the azimuth angle\u03c8 which is measured from the downstream position of the blade. Being pivoted to the main rotor shaft, the Bell\u2013 Hiller stabilizing bar is a steel stick equipped with two paddles. Tilt of the swashplate, as is illustrated in Fig. 3, directly maneuvers the cyclic pitch of the paddles, which in advance contributes to a cyclic flapping motion of the Bell\u2013Hiller stabilizing bar. Since the stabilizing bar is a single rigid body, no collective flapping exists. The flapping motion of the stabilizing bar can change the cyclic pitch of the blades, which is shown in Fig. 4. The above relationships can be expressed by the following equations: B f = \u03b4 \u00b7 l\u03b4l2 l\u03b2 f l1 , (1) \u03b2\u0308 f = f f (\u03b2\u0307 f , \u03b2 f , B f ), (2) \u03b8 = \u03b2 f \u00b7 l3l5 l4l\u03b8 + \u03b4 \u00b7 l\u03b4(l5 \u2212 l4) l4l\u03b8 , (3) \u03b2 = fb(\u03b2\u0307, \u03b2, \u03b8), (4) where (1) and (3) are explained in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003694_ecce44975.2020.9235404-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003694_ecce44975.2020.9235404-Figure4-1.png", + "caption": "Fig. 4. Two configurations of the integrated motor-compressor, (a) aligned stator and rotor with 0 skew stator-rotor skew angle, (b) unaligned stator and rotor with stator-rotor skew angles.", + "texts": [ + "4 Axial flow velocity, V1 [m/s] 140 Mass rate of flow [kg/s] 0.664 Rotor thermodynamic efficiency 98.3% Stage thermodynamic efficiency 90.5% Ideal total pressure ratio 1.088 Current density [Arms/mm2] 15 Number of turns per coil 7 Rated current [Apk] 54 Motor output torque [Nm] 3.86 Motor efficiency 92.2% The airfoil-shaped rotor has a stagger angle. Normally, in order to maximize the torque production capability of the integrated motor-compressor, the stator and magnets need to be skewed by the same stagger angle to align with the rotor in Fig. 4(a), where the stator skew angle is equal to the rotor stagger angle [8]. In this paper, the unaligned stator-rotor configurations in Fig. 4(b) are evaluated. In such configurations, the airfoilshaped rotor is skewed by the stagger angle. However, the stator and magnet are not aligned with the rotor, and the stator skew angle is not equal to the rotor stagger angle. In order to account for the effect of skew in the stator and magnets, continuous-skew and step-skew are applied to the stator and magnets, which are shown in Fig. 1, and Fig. 4(a), respectively. The manufacturing of the step-skewed structure for 5546 Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 22:42:22 UTC from IEEE Xplore. Restrictions apply. the stator and magnet is much easier to realize than the continuous-skewed structure. Fig. 5 shows the front view and top view of the stator of an integrated motor-compressor. In Fig. 5 (a), the circumferential skew angle of the stator \u03b1s is calculated in (1) where r is the stator inner radius, \u03b3 is the rotor stagger angle or stator skew angle, x is the approximate skewed pitch at stator inner radius, and le is the stack length of the stator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure4-1.png", + "caption": "Figure 4: Free-state stiffness measurement system, assembled rig photo", + "texts": [ + " Using the axial spacing of 25% and tangential spacing value of 5%, the number of bristle rows in seal axial direction is found to be 15 by using the formula for a staggered bristle layout given by Aksit [7] (also visualized in Figure 3). 2 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use A free-state stiffness measurement system has been designed to characterize the seal stiffness under unpressurized static rotor conditions. The measurement relies on pressing a metallic pad, which has an arc surface of the same diameter as the seal inner diameter, against unloaded bristles. As detailed in Figure 4, the metallic pad movement is generated by the linear slide. A controlled input is given to the slide by using the controller and a PC. As the linear slide is moved, the load cell measures the force exerted to the metallic pad by the bristles. At the same time, the displacement sensor measures the magnitude of metallic pad \u2013 seal interference. The load cell and displacement sensor are connected to a data acquisition system to collect and store the data on a computer. An amplifier is used to increase the quality of the recorded test data and the accuracy of measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001639_urai.2015.7358875-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001639_urai.2015.7358875-Figure2-1.png", + "caption": "Fig. 2: Configuration and workspace of TRS", + "texts": [ + " \\Vheels are attached to the arms which can be folded and unfolded by air pressure, which provides adhe sive force of the wheels to the pipe inner surface and also damping to the sudden motion of the arm due to external disturbance to the wheels. The traction force would be optimal when the entire bodies of the robot aligns with the pipe axis if straight. So, our path planning aims to keep the driving modules on the axis of the straight pipe. In addtion, there should be at least two tractable driving modules when the robot passes through a non-straight pipe such as an elbow pipe. It is noted that the roll motion of the robot which seldom occurs is ignored in this paper. The configuration of TRS is depicted in Fig. 2a and the workspace representation of the robot is shown in Fig. 2b. Notice that the location of the joint i is represented by (Xi+l' Yi+l) in workspace and its cor responding joint angle is denoted by (Pi- 2.2 Configuration Space and Workspace RRT extends the node of tree, which is nearest to a random sample and 'nearest' refers to a distance metric between nodes in configuration space. How ever, in case of inpipe robots, the 'nearest' node may not correspond to the configuraiton that we expect. For example, in Fig.3, the configuration distance be tween nodeo and nodea is shorter than that between nodeo and nodeb because the former has only one joint angle is non-zero but the later has two" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000630_ilt-11-2011-0103-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000630_ilt-11-2011-0103-Figure3-1.png", + "caption": "Figure 3 Barrier design", + "texts": [ + " In such a case, the vertical clearance of both the bush halves is the same [cv1B cv2B cc, in agreement with equations (6) and (7)], as the edge of the top pad is tangent to the cylindrical profile machined by the first turning operation (Figure 2 with tv2 0). Independently of the design solution, with reference to Figure 2, the cut depth at the bottom and top pad center (on the load line), tv1 and tv2, respectively, in agreement with equations (2) and (7) are defined by: tv1 e1 t (8) tv2 cv2 cc t e2 (9) The second turning operation may be performed either on the whole length L of the bearing land (in the axial direction) or on a part of it, symmetrically with reference to the load line. In the latter case, schematized in Figure 3, on each atmospheric side of the bearing, a cylindrical surface with axial length Lb and uniform radial clearance cb cc is left. As cc is the smaller clearance of the journal housing, the two \u201cbarrier\u201d surfaces should reduce the lubricant side loss effectively also in nominal working conditions. In summary, the bearing (out of the barrier zones) is made up by three pads: two eccentric pads of the \u201celliptical\u201d part (with eccentricity e1, e2, respectively, and clearance c1 c2 cp) and a pocket pad (with eccentricity e3 0 and clearance c3 cc )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003112_0954407020916991-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003112_0954407020916991-Figure7-1.png", + "caption": "Figure 7. The final finite-element model of the single sun gear tooth.", + "texts": [ + " So a layer of SURF 152 thermal effect units need to be set on the working surface to play a role as heat flux density. The final finiteelement model of the single sun gear tooth is shown in Figure 4. The experimental setup of the coupling system. Table 3. The temperature of some points in thermal network method and experimentation. Point 1 Point 2 Point 3 Point 4 Point 5 Temperature in thermal network method ( C) 45.9 39.5 38.0 45.7 52.4 Temperature in experimentation ( C) 46.6 34.9 35.7 51.2 51.7 Figure 6. The model of the single sun gear tooth. Figure 7, and the result of the FEA is shown in Figure 8. As shown in Figure 8, the temperature of the working surface is obviously higher than the temperature of other parts. The temperature distribution from the middle of the tooth surface to the end face is graded distribution, and it is completely symmetrical throughout the tooth width. These temperature distributions are in line with the actual situation. The temperature field distribution on the tooth profile line right in the middle of the meshing tooth surface was extracted from ANSYS, and the data were imported into MATLAB for analysis, the results are shown in Figure 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002017_1.4033364-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002017_1.4033364-Figure15-1.png", + "caption": "Fig. 15 Bearing modeling in ANSYS", + "texts": [ + " Again, as the fastening screws connecting the test table and the bearing inner ring are closely distributed to the mounting area, it is a good approximation to fix the bearing from the lower face omitting fasteners in the model. Axial load and moment are applied on the upper face of the outer ring. Standard earth gravity is defined in \u201c y\u201d direction. Finally, a cylindrical support is defined to restrict the rotation of outer ring with respect to the inner ring in \u201cy\u201d direction. Representations of boundary conditions and loadings are given in Fig. 15. 4.2 Roller Modeling in ANSYS. Modeling the bearing with all the details causes too much solution time, since it has many elements and contacts between them. The most critical part to model for the load distribution analysis is the rollers. Rollers make contact with inner and outer rings transferring the force between them. So, in the finite element analysis program, rollers are modeled as compression only springs. Spring constants are calculated as given in Ref. [15] Fn \u00bc K1dm (3) where K1;x \u00bc plE0 2 2 3 \u00fe ln 2d1;x bx \u00fe ln 2d2 bx (4) dm \u00bc dmo \u00fe dmi (5) The x-subscripts in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001380_amm.554.515-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001380_amm.554.515-Figure1-1.png", + "caption": "Fig. 1 Seven degree of freedom of vehicle ride model", + "texts": [ + " In this study, a ride model is derived based on Ikenaga (2000) and later being validated by the experimental data. The validation on the developed model is necessary in order to make sure that the developed model is valid to be used in order to further study vehicle\u2019s ride dynamics or to study advance suspension system. A vehicle\u2019s ride model is derived based on the work done in [5]. The ride model consists of seven degrees of freedom namely roll, pitch, bounce and vertical motion of each four wheels. Figure 1 show the vehicle\u2019 ride model. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 141.211.4.224, University of Michigan Library, Media Union Library, Ann Arbor, USA-18/04/15,23:29:46) Based on the 7DOF of ride model in Figure 1, the displacements of the sprung masses are given by; cariLcar cara bZsijZ \u03b1\u03b8 \u2212+= 2 (1) with sijZ is the total sprung mass displacement (i =f for front, r for rear and j=l for left, r for right), bZ is the sprung mass vertical displacement at the center of gravity, car\u03b8 is the roll angle and car\u03b1 is the pitch angle. The distance of centre of gravity to the front axle and rear axle are given by fL and rL respectively. The forces acting at each of the suspension )( ijF is the sum of the spring force )( sijF and damper force )( dijF " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.21-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.21-1.png", + "caption": "Figure 10.21 The (a) lower-tip-point and (b) higher-tip-point posture modes of the manipulator.", + "texts": [ + " According to Eq. (10.528), such a pose can occur for the legs L2 and L3, i.e. for k = 2 and 3, if \ud835\udf0ek = sgn(nk3) (10.532) As for the first leg, Eq. (10.521) implies that such a pose can occur if \ud835\udf0e1 = sgn(K1) (10.533) However, as verified in Section 10.8.7, the moving platform remains always parallel to the fixed platform. Therefore, the lower-tip-point and higher-tip-point posture modes of the manipulator must be realized consistently by all the three legs. These two posture modes are illustrated in Figure 10.21, in which only the first and third legs of the manipulator are taken into the view as if they are coplanar (i.e. as if \ud835\udefd3 = \u2212\ud835\udf0b) for the sake of neatness. Here, it may be needless to say that the delta robot is of course preferred to be in the lower-tip-point posture mode in its customary operations. (c) Position Singularities of Forward Kinematics The forward kinematic solution presented in Part (a) stipulates three conditions in order to obtain a definite pose of the manipulator corresponding to a specified set of active joint variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000788_metroaerospace.2014.6865903-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000788_metroaerospace.2014.6865903-Figure5-1.png", + "caption": "Fig. 5. Case study: snapshot (top), voxel structure (down).", + "texts": [ + " Percentile analysis gives a deeper overview of the distance distribution since it computes the averaged distance as the distance between the median position (taken from the best fitting planes) of each face and the maximum and minimum values as percentile distances (10% or 90%). This allows to cut off outliers or couples of points that are not statistical relevant for the measurements. III. CASE STUDY The proposed application aims to apply the procedure to an aeronatical flange shown in Fig. 4 and Fig. 5. It is a component of a Boeing product, obtained through a \u03b2-forging process starting from Ti6Al4V Titanium Alloy Fig. 3. Example of dimensional and planarity inspection of two planar surface. powders, that is being studied by Centro Sviluppo Materiali. This powder material is suited for manufacturing near net shape forged components since it has been demonstrated that it is possible to manufacture by powder-metallurgy samples that are mechanically equivalent to those from commercial extruded bar [10, 11]", + " The laser scanner has a probing error of 2.5 \u03bcm (obtained with test comparable to EN/ISO 10360-2) and can acquire approximately 70,000 points/s, and 900 points per line. The CMM system has a range of accuracy of about 2\u00f76 mm with a positioning maximum speed of 750 mm/s. To avoid reflections the surface has been white-mat painted. Doing so, after a multiple view recomposition, a cloud of 479780 points has been achieved and analysed according to a voxel structure of 64x64x64 voxels, as shown in Fig. 5. Following the segmentation scheme described in section 2, planar surfaces and cylinders are recognized. Fig. 6 shows the voxel\u2019s values of the standard deviations from the local best fitting plane, taken in a transversal section of the flange. Fig. 6 shows that planar surfaces always have standard deviation smaller then the cilindric ones, with a ratio of about 1:2. This allows to find a significant threshold for distinguishing cylinders from planes. The IRS is found starting from the direction orthogonal to the transversal section of Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001463_iciea.2015.7334434-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001463_iciea.2015.7334434-Figure1-1.png", + "caption": "Fig. 1. The stator winding structure of DYSP-PMSM.", + "texts": [ + " Among all multi-phase motors, five-phase AC motor and dual Y shift 30 degree six-phase permanent-magnet synchronous motor (DYSP-PMSM) is investigated most extensively. The stator winding of DYSP-PMSM is consist of two sets of three-phase winding, there are many welldeveloped theories and experience can be borrowed from three-phase PMSM, so DYSP-PMSM is chosen as the study object. The biggest advantage of direct torque control (DTC) is rapid response of control system. DTC of multi-phase PMSM, which has higher precision and smaller torque ripple than three-phase PMSM, is necessary to be studied further. The stator winding structure of DYSP-PMSM is shown in Fig.1. The stator winding is consist of two sets of threephase winding, which are labelled as ABC and XYZ, differ by 30 degree from each other, and star connection method with isolated neutral points is used. In natural coordinates, analysis of DYSP-PMSM is complicated because it is a multi-variable, nonlinear and strong-coupling system. According to principle of motor model equivalence, the magnetic potentials resulting from stator windings of motor are equal in different coordinates [5]. Make two transformations with motor, from six-phase asymmetric windings A, B, C, X, Y, Z to two-phase static windings ,\u03b1 \u03b2 (named 6s/2sC transformation) and from twophase static windings ,\u03b1 \u03b2 to two-phase rotated windings ,d q (named 2s/2rC transformation), two transformation matrices are expressed as following: 6s/2s 1 cos(120 ) cos(240 ) cos(30 ) cos(150 ) cos(270 ) 0 sin(120 ) sin(240 ) sin(30 ) sin(150 ) sin(270 ) 1 1 cos(240 ) cos(120 ) cos(150 ) cos(30 ) cos(270 ) 3 0 sin(240 ) sin(120 ) sin(150 ) sin(30 ) sin(270 ) 1 1 1 0 0 0 0 0 0 1 1 =C 1 (1) 4 2s 2r 4 4 4 4 cos( ) sin( ) 0 sin( ) cos( ) 0 T T \u03b8 \u03b8 \u03b8 \u03b8= \u2212C 0 0 I (2) Where \u03b8 is the angle between d axis and \u03b1 axis, 40 is row vector and 4I is a four-dimensional unit matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003796_icem49940.2020.9270813-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003796_icem49940.2020.9270813-Figure2-1.png", + "caption": "Fig. 2: Geometric depiction of the stators with the excitation points (11-38) and the measurement point (21).", + "texts": [ + " When using the moving sensor method, the body is always excited at the same position and the sensor is moved around at different positions on the body. The FRFs of the different points are measured and saved. The peaks in the FRFs indicate eigenpairs of the body. The corresponding eigenshapes can be animated and identified. To carry out a modal analysis, to animate the eigenshapes and to determine the modal damping, different software approaches exist. In this work, the M+P Analyzer is used to perform the modal analysis on the stators. The geometry of the stator core is approximated with 24 points, as it can be seen in figure 2. Eight points at three axial levels are set and named accordingly (11\u2212 18, 21\u2212 28, 31\u2212 38). The acceleration sensor is placed at position 21. This set-up allows the identification of radial eigenmodes of the order \u03c1 < 3. The eigenmode \u03c1 = 4 can be identified as well, since the system is excited at the position of the sensor and thus the maximum amplitude is measured. Higher eigenmodes \u03c1 > 4 appear as eigenmodes of lower order \u03c1 < 4 according to the Nyquist criterion. To perform the modal analysis, the stator cores are hung on an elastic band that is fixed to a frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.26-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.26-1.png", + "caption": "Figure 9.26 First kind of position singularity of a Scara manipulator.", + "texts": [ + "299) also implies that this singularity occurs if the position of the end-effector is specified in such a way that the point Q coincides with the base frame origin O. That is, r1 = r2 = 0 (9.313) The above equations express the noticeable feature of this singularity that the upper and front arms of the manipulator have equal lengths and the front arm is folded completely over the upper arm. However, this singularity cannot occur exactly because of the physical shapes of the relevant links and joints. This singularity is illustrated (approximately) in Figure 9.26. If this singularity occurs, \ud835\udf031 becomes indefinite and ineffective. In other words, \ud835\udf031 can be assigned an arbitrary value, but it cannot cause any change in the position of the end-effector, whatever the assigned value is. Thus, the manipulator gains a positioning freedom in the joint space, but this freedom happens to be useless in the task space as long as the points Q and O are kept coincident (actually almost coincident). On the other hand, if the task to be executed necessitates to keep the manipulator in such a pose, then the freedom in \ud835\udf031 may be used to orient the folded arm links of the manipulator conveniently depending on the environmental conditions, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003949_j.jterra.2020.11.004-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003949_j.jterra.2020.11.004-Figure4-1.png", + "caption": "Fig. 4. The experimental setup for the wheel towing test.", + "texts": [ + " For calculation F5, the shear stress s\u00f0h\u00de in the tangential direction of the wheel surface at an arbitrary point on the wheel surface is given as follows: s\u00f0h\u00de \u00bc \u00f0cw \u00fe r\u00f0h\u00de tan d\u00de\u00f01 e j=j\u00de \u00f012\u00de where cw is the soil-wheel adhesion, d is the tool-soil friction angle, h indicates the wheel contact angle, and r\u00f0h\u00de indicates the normal stress. Therefore, the sum of the horizontal component of the shear force s\u00f0h\u00de along the wheel surface is the force F5 between the wheel surface and soil as follows: F5 \u00bc wr Z ho 0 s\u00f0h\u00de cos\u00f0h\u00dedh \u00f013\u00de where w is the wheel width and r is the wheel radius, and ho indicates the wheel contact angle at the soil surface. The calculation is performed using the soil and wheel parameters Tables 2 and 3. Fig. 4 shows the schematic and overview of the experimental setup for the towing test. The size of the wheel is set based on the rover testbed size (wheel diameter, width, and the total mass are 170 mm, 40 mm, and 10 kg) owned by our laboratory (Fujiwara et al., 2019). To confirm resistance force at each size, the size of the wheel is set at 170 and 350 mm in diameter, 40 Table 2 Soil parameters and values. Modulus Value Unit Name of parameters Reference c 762a \u00f0N=m2\u00de Soil cohesion Mizukami, 2013 C0 same as c \u00f0N=m2\u00de Soil-tool adhesion for a virtual plate Mizukami, 2013 cw 0 \u00f0N=m2\u00de Soil-wheel adhesion \u2013 g 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.7-1.png", + "caption": "Figure 9.7 Second kind of position singularity of a Puma manipulator.", + "texts": [ + "48) also implies that this singularity occurs if the position of the end-effector is specified in such a way that the wrist point coincides with the shoulder point. That is, r12 = \u221a r2 1 + r2 2 = d2 (9.64) r3 = 0 (9.65) The above equations express the noticeable feature of this singularity that the upper and front arms of the manipulator have equal lengths and the front arm is folded completely over the upper arm. However, this singularity cannot occur exactly because of the physical shapes of the relevant links and joints. This singularity is illustrated (approximately) in Figure 9.7. If this singularity occurs, \ud835\udf032 becomes indefinite and ineffective. In other words, \ud835\udf032 can be assigned an arbitrary value, but it cannot cause any change in the position of the end-effector, whatever the assigned value is. Thus, the manipulator gains a positioning freedom in the joint space, but this freedom happens to be useless in the task space as long as the wrist point is kept coincident (actually almost coincident) with the shoulder point. On the other hand, if the task to be executed necessitates keeping the wrist point as such, then the freedom in \ud835\udf032 may be used to orient the folded arm links of the manipulator conveniently depending on the environmental conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001603_j.proeng.2015.11.237-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001603_j.proeng.2015.11.237-Figure2-1.png", + "caption": "Fig 2 Structure and Gas film action layout of gas lubrication self-acting bearing", + "texts": [ + " So as to realize the instrument with high accuracy long time continuous work and long life requirements. This paper counter complex flow characteristics clearance of cone self-acting gas lubrication bearing, using the computational fluid dynamics(CFD) numerical simulation method to carry out the research on the static characteristics of the bearing, and the results were compared with the experimental results. Cone self-acting gas lubrication bearing structure and lubricating film force schematic was shown in Figure 2, the structure was divided into two parts the rotor and the stator. The stator was composed of two groups of cone on the composition, supported the rotor. When the bearings worked in high speed rotation, bearing clearance, the rotor cone wall would drag outside lubrication gas from the wide passage to the narrow channel depended on the gas viscous effect.As a result, there is obstruction, which would make the pressure rise to become high-pressure area, wide channel gave birth to the low pressure area, which is the basic principle of Cone self-acting gas lubrication bearing that produces load ability, also known as the air bearing dynamic pressure effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000903_tmag.2014.2359575-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000903_tmag.2014.2359575-Figure1-1.png", + "caption": "Fig. 1. Search coil locations in the measurement sample.", + "texts": [ + "2014.2359575 size holes of non-oriented fully processed electrical steel sheet. A non-oriented 300 mm long, 30 mm wide, and 0.5 mm thick steel sheet, which was namely N530, was drilled with 10 and 20 mm diameter to make a hole at the center. The 1 mm holes, which are minimizing the stress, were drilled through the sample at proper angles to wound 10-turn search coils. The four search coils for each hole with 0\u00b0, 25\u00b0, 45\u00b0, and 65\u00b0 angles corresponding to the center of hole were located as shown Fig. 1. A 100-turn 0.1 mm enameled copper wire main search coil (MC) was also wound uniformly around the sample to detect the overall peak flux density. The drilled sharp edges are isolated to avoid any contact the search coil wires with the sample. These samples were magnetized sinusoidally and parallel to the cut edge over the frequency 50\u2013400 Hz for flux density level up to 0.5 T to avoid flux instabilities in a single-sheet tester. The single-sheet tester comprises the strip magnetizing and secondary windings and U-shaped flux return core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001437_s0263574714001398-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001437_s0263574714001398-Figure7-1.png", + "caption": "Fig. 7. (Colour online) (a) Free body diagram of rigid parts of each robot leg; (b) Prismatic joint effects on the beam.", + "texts": [ + " Amount of the slope angle of deformation at points xpi is called \u03b8pi . To apply the effect of prismatic joint length Lp following constraints can be written using Lagrange\u2019s multipliers FR pi and FL pi as FR pi ( \u03b4wi ( xpi + Lp/2 ) \u2212 Lp 2 \u03b4\u03b8pi ) = 0, i = 1, 2, 3, FL pi ( \u03b4wi ( xpi \u2212 Lp/2 ) + Lp 2 \u03b4\u03b8pi ) = 0, i = 1, 2, 3, (28) in which the Lagrange\u2019s multipliers FR pi and FL pi represent normal surface forces that are applied to the beam by the prismatic joint at its right and left edges, respectively (Fig. 7(b)). Next, the effect of rotational inertia of the prismatic joints in the dynamic model is considered as \u03b8pi = \u2202wi(xi, t) \u2202xi \u2223\u2223\u2223\u2223 xi=xpi = \u2202wi \u2202xi (xpi, t) \u21d2 Mpi ( \u03b4\u03b8pi \u2212 \u03b4 \u2202wi \u2202xi (xpi) ) = 0, i = 1, 2, 3, (29) in which Mpi is a Lagrange\u2019s multiplier and represents the moment that the beam applies on the corresponding prismatic joint. According to constraint Eq. (24), the axial displacements of beams xiG are included in the dynamic model. Then, to solve for xiG, motion equations for axial motion of the beams are obtained by considering the virtual work principle, Eqs", + " (30)\u2013(32), must be simultaneously solved with the constraint equations, Eqs. (24)\u2013(29). Upon solving these equations, parameters Fpi and \u03b8pi are determined. In addition, by specifying mass of the prismatic joints, required driving forces, Fai , can be determined as follows: (mpp + mpa)si (t) = Fai \u2212 ( Fpi + FL pi + FR pi ) cos ( \u03b8 + \u03b8pi ) , i = 1, 2, 3, (33) in which parameters mpp and mpa represent masses for the passive and the active prismatic joints, respectively. Note that the two prismatic joints are joined by a revolute joint (see Fig. 7(a)). Upon obtaining Fai , the inverse dynamics problem of the ST robot is solved. It should be noted that the same formulation allows solving for the direct dynamics problem of the robot. This may be accomplished by specifying the driving forces, Fai , in Eq. (33) and solving a new set of motion equations, Eqs. (30)\u2013(33) simultaneously, with the constraint Eqs. (24)\u2013(29). In this section, the \u201cconstrained assumed modes method\u201d is further developed to solve the derived motion equations of the ST parallel robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002017_1.4033364-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002017_1.4033364-Figure1-1.png", + "caption": "Fig. 1 Three wire-race double row roller slewing bearing", + "texts": [ + " Friction torque measurement results are investigated and compared with analysis results. Friction sources are decoupled and examined separately in order to fully understand effect of each component on friction torque of the slewing bearing. Moreover, a valid friction torque estimation of the slewing bearing under different load conditions can be simulated without the need of tests in the design stage [1]. The type of the slewing bearing being investigated in this paper is a three wire-race double row roller bearing with a mean diameter of 552.2 mm, purchased from Tibet Makina1 (Fig. 1). Wirerace bearing is a special type of slewing bearing, that is, it is made lighter by using aluminum inner and outer rings instead of steel. However, for the load to be supported, the race is composed of steel wires. Wire-race slewing bearing is mainly composed of six types of parts: inner ring, outer ring, rolling elements, raceways, seals, and cages. Relative motion occurs between inner and outer rings while rolling elements roll between wire races. Cage keeps rolling elements in equidistant positions, hence prevents any collision between rolling elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002483_itec-india.2015.7386862-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002483_itec-india.2015.7386862-Figure13-1.png", + "caption": "Figure 13. Fuel Cell Arrangement - 1", + "texts": [], + "surrounding_texts": [ + "4. Quick position change ability, as jumping out of cover position, stealth mode, and underwater operation without Diesel engine running; 5. Energy saving and regeneration e.g. during braking operation; 6. Electric energy supply to all consumers during standstill of the vehicle, e.g. silent watch; 7. Option for future internal and external supply of high power consumers. 8. According to the wide area of the different tasks greater energy and greater power have to be provided. 9. In addition short duration and long duration of use have to be taken into account. 10. High level of energy is needed for long time application to meet the power requirement of the subsystems e.g. for system initiation/activation or silent watch. High level of power is needed for: The start of the prime mover, Mobility e.g. acceleration, 11. Weapon power supply like ETC-gun, Active armour, Active suspension.\nThe magneto-dynamic storage is a flywheel storage with an integrated electric machine that can be used either as a generator (discharge mode), or as a motor (recharge mode via re-accelerating the flywheel rotor) depending on the momentary needs. The energy carrier of the MDS is a cylindrical rotor made of wound carbon fiber. The rotor\u2019s axle stands on a vertical plane. The motor/generator (M/G) unit is inside the cylindrical rotor, accepting or delivering electric power. To reduce the friction in the bearings most of the rotor\u2019s weight is compensated for by magnetic forces. Air friction is reduced to a minimum as the complete rotor unit runs in a vacuum enclosure.\nThe most important advantages of the MDS compared to other energy storages, i.e. chemical batteries are first of all the high power ability with respect to weight and volume and the indefinite cycle number potential. This is due to the fact that the MDS is an electric machine and is not limited by electrochemical elements. These characteristics open the benefits to use the MDS in military vehicles as an additional energy and power source. The results are the excellent features of the MDS and its related advantages.\nPresently the MDSs reach specific values of energy/mass = 80 MJ/ton and power/mass = 6.5 MW/ton in the laboratory. In the next few years approximately 15 MW/ton will be achievable.\nThis MDS type has been safe proved by the international technical authorities and is authorized for use in public transport applications. With more than 200,000 operation hours and more than 2.5 million kilometres in vehicles this technology stands for an evident reference.\nFuel cells generate electrical energy by an electrochemical reaction (i.e. without combustion). As a result, fuel cells offer a high potential efficiency, and emit exhaust gases comprised solely of water vapour. In contrast to the majority of prime movers, the efficiency of fuel cells is the greatest at partial loads rather than at full load.\nConsequently, other than is the case with internal combustion engines, it is not beneficial to down-size a fuel cell solely on the basis of efficiency considerations \u2013 for example, as a component of a series hybrid system (a relatively small fuel cell and a constant high load, instead of a large fuel cell usually subjected to a partial load). However, since the available fuel cells are still heavy and bulky \u2013 and, above all, expensive \u2013 many projects will nevertheless deploy a small fuel cell. The expected reduction in weight, volume and price will lead to a trend towards the use of fuel cells as a component of the power train. In view of the use of electricity produced by fuel cells to power vehicles the recuperation of the braking energy will continue to be of interest; this can be achieved by the use of a small battery to serve as a buffer, and consequently these systems will continue to be of a hybrid nature. The inherent inertia of reformers may result in the need for Systems employing a reformer to make use of a hybrid power train so as to provide for a rapid response to the operation of the gas pedal. Fuel cells are often regarded as a potential longer-term (several decades) serious alternative to the internal combustion engine.\n1. High potential efficiency, 2. Low emissions, 3. Opportunity to use hydrogen produced from sustainable\nsources, 4. Possibly less maintenance as a result of the absence of\nmoving parts, 5. The opportunity to make use of the existing fuel\ninfrastructure in combination with reformers, 6. Hydrogen is extremely inflammable, and difficult to store, 7. The complete absence of a hydrogen infrastructure for\nsupplies to ordinary vehicles, 8. High costs, 9. The use of a reformer lowers the efficiency, 10. Slow response of the reformer to varying loads, 11. The power train can be of a substantial size,\nSystem integration.\nFuel cells convert chemical energy directly into electrical energy. This process does not involve the Carnot cycle, and consequently extremely high electrical efficiencies (to 80%) are theoretically possible. The fundamental difference between fuel cells and batteries pertains to the fuel and the oxidant. In batteries the fuel and oxidant are stored in the form of a solid or a liquid in the battery (and are regenerated when the battery is charged), whilst fuel cells receive continual supplies of fuel and oxidant in the form of a gas (and sometimes as a liquid) from an external source. Consequently, in analogy with combustion engines, the range of vehicles propelled by fuel cells is determined by the content of the (separate) fuel tank.\n\u2022 ALKALINE FUEL CELLS (AFC) \u2022 SOLID OXIDE FUEL CELLS (SOFC) \u2022 PROTON EXCHANGE MEMBRANE \u2022 FUEL CELLS (PEMFC) \u2022 PHOSPORIC ACID FUEL CELLS (PAFC) \u2022 MOLTEN CARBONATE FUEL CELLS (MCFC)", + "Experiment Tank & Notional Tank - 2 required fuel cells\n\u2022\n\u2022 \u2022\n\u2022\n\u2022\n\u2022\n\u2022\n\u2022\n\u2022\nUse of Fuel Cells have been already incorporated by many research organizations around the globe, which makes it all the more feasible and workable with as some background research material is available. Some more advantages of Fuel Cell Research:\n\u2022 The reduction of the response time from >10 min. to < 1 min; \u2022 The increase of the specific power (W/kg) by means of the\nintegration of the components; \u2022 The reduction of the costs; \u2022 The improvement of the efficiency. \u2022 It is evident that the use of hydrogen as the fuel will require\n(and certainly during the coming ten to twenty years) the deployment of the centralized reforming of (fossil) fuels (for example, as service stations).\n\u2022 This reforming process can be affected in large installations in which issues such as the response time, compactness, etc., are much less critical, and in which higher efficiencies can be much more readily attained.\nENERGY STORAGE" + ] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure8-1.png", + "caption": "Figure 8. Condition of constraints of a unit.", + "texts": [ + " To this end, the coordinates of the joints are solved by kinematics: (3) Meanwhile, Pi (x, y, z) is the coordinate of a joint, Ti is a homogeneous transformation matrix, and N is the number of units in the robot. Meanwhile, the coordinates of P0 and P1 are previously defined. We define the condition of the posture of the robot to adsorb onto spherical surfaces. First, we calculate the condition of constraint when both of the two rollers of a magnetic adhesion mechanism adsorb onto a spherical surface, as shown in Fig. 8. Then, the distance between each joint and the center of the sphere is given by (4). (4) Here, rc is the vector from the center of the sphere to the center of the unit, lc is the vector from the center of the unit to its tip, dc is the vector from the center of the sphere to the midpoint between the joint and the other adjacent unit, Lhi-1 is the vector from the midpoint to the joint and d is the distance between the joint and the center of the sphere. We consider that the value of d is constant whether the robot is moving or fixed on the surface", + " \u2212 + + \u2212 \u2212 = ii iiiii iiiii ii iiiii iiiii i ii ii i c\u03c8\u03b8c s\u03c8cc\u03c8\u03b8ss s\u03c8sc\u03c8\u03b8sc s\u03c8\u03b8c c\u03c8c\u03c8s\u03b8ss c\u03c8ss\u03c8\u03b8sc \u03b8s \u03b8cs c\u03b8c \u03c6\u03c6 \u03c6\u03c6 \u03c6\u03c6 \u03c6\u03c6 \u03c6 \u03c6 R = \u2212 0 0 1i i i L P , 1 , 10 , 1 1 =\u2032 = = \u2032= \u2212 i i i i ii i i i i i ii P P PR T P P PTP r r rr 1,,3,2 \u2212= Ni L ccc hici d lrd LdP rrr rr += +== \u22121 Next, we define the condition of the posture of the robot in the case of moving and being fixed. To define the condition, we reflect the posture of each unit in the modeled robot, as shown Fig. 8. Then, Pui is the coordinate of the center of the unit, and Pri is the coordinate of the roller B. Pui and Pri are given by (5) and (6), respectively. (5) (6) Here, l is the length between the center of a unit and its tip, and lr is the length between the center of the roller B and the tip of a unit. When the robot is fixed, the vector normal to the unit N, which passes through Pui, crosses the center of the sphere, as shown Fig. 10 (a). Then, Pui satisfies (7). (7) When the robot is moving, both of the rollers of the magnetic adhesion mechanism cannot adsorb mechanistically onto spherical surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002121_tcst.2016.2572165-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002121_tcst.2016.2572165-Figure8-1.png", + "caption": "Fig. 8. Sources of guidance actuator force errors. (a) Possible misalignment due to angle of attack. (b) Misalignment within the plane of action.", + "texts": [ + " Finally, we may increase \u03bbg for faster convergence to arbitrarily reduce the error and as a result d approaches an arbitrarily small value. Corollary 2: The sliding-mode control law may be replaced with a finite time controller [25], such that the target is reached precisely. However, in practice, more complex control laws are unnecessary due to many sources of error as described in Section IV-A. The first source of error is in possible misalignment outside of the plane of action due to a nonzero angle of attack, as shown in Fig. 8(a). Noting that we are considering impulsive forces, the control force, Fguid, derived in (21) must be available in the direction of desired change in velocity, vdiv, to maintain an exponentially stable guidance velocity error. Assuming a nonzero angle of attack, the desired guidance control force direction may have a component in the body x-axis. However, the guidance force provided by the thrusters can only be provided within the yz plane. We denote the angle between the udiv-direction and the yz plane, by the tilt angle \u03c4 \u03c4 = sin(i \u00b7 udiv), 0 \u2264 \u03c4 \u2264 \u03b1", + " (26) The direction of vdiv and v \u2032 div are given as \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 udiv = vdiv ||vdiv|| = [sin \u03c4 \u2212 cos \u03c4 sin \u03b2 cos \u03c4 cos\u03b2]T u\u2032 div = v \u2032 div ||v \u2032 div|| = [0 \u2212 sin \u03b2 cos\u03b2]T . (27) It is important to note that this error can be avoided by a good attitude control law that maintains a zero angle of attack. A second source of guidance control error is due to in-plane misalignment. This is because u\u2032 div will not normally be aligned with the nearest guidance thruster located at angle \u03b2i relative to the body z-axis, as shown in of Fig. 8(b). Therefore the actual guidance control force will be applied along ua div = (va div/||va div||) = [0 \u2212 sin \u03b2i cos\u03b2i ]T , and the magnitude of actual contribution of velocity change in the desired direction will be ||va div|| = ||vdiv|| cos \u03c4 cos(\u03b2 \u2212 \u03b2i ). (28) This error may be reduced by providing an initial roll rate such that the thrusters become aligned with the desired direction after a short wait time. In addition, the more the thrusters we use, the smaller this error will be. The main objective of attitude control is to maintain a zero angle of attack throughout the flight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002755_1.5063139-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002755_1.5063139-Figure8-1.png", + "caption": "Figure 8: Temperature contour with red representing regions above melting point; all pictures created at 7 load steps past when material was first. Left \u2013 r0.5 len10 Middle \u2013 r1 len10 Right r1 len15", + "texts": [ + " Blue \u2013 Green r1 as observe ceeds the m an 5\u00b0C it is not reflec it does emu ntly to predi erature at star RAD1 307.9 350.84 389.85 424.36 455.04 420.4 448.3 448.11 ble 3 (Fig ap layer inc aterial are a ature (top) an r from the load r0.5 len10 Re len15 d that if a elting point considered t the actual late the bulk ct failure. t of cap RAD 2 341.76 400.14 451.39 495.68 534.21 568.17 598.5 621.11 ure 7) the reases until dded. d heat flux step where d \u2013 r1 len10 While only the BCC of cubic Size 15 is shown in Figure 8, all the 15mm cubics with 1mm radius struts failed in the same way (see Figure 9). There are significant future developments that could be made to the simulation. For example a process known as rezoning will allow for the generation of a higher density mesh only near where the heat flux is applied. Further verification of the model is required to ensure results are consistent and identify all relevant failure modes. For example, the failure of BCC part with radius 1 and cubic size of 10mm was unexpected as the FCZ and FBCZ of the same size both were moderately manufacturable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003416_aim43001.2020.9158920-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003416_aim43001.2020.9158920-Figure2-1.png", + "caption": "Fig. 2: Computer aided design illustrations of the system.", + "texts": [ + " A plastic holder was designed to integrate the two reaction wheels and to mount them onto the link. The flywheels were fabricated from a steel sheet of 3 mm thickness using a laser cutting machine (Bodor P3015). The mass of each flywheel is 160 grams and its radius is 8 cm, and their design was aimed at maximizing the moment of inertia. The flywheels are covered with external plastic cases for safety, and their angular speeds are measured by two encoders. Finally, two gyro sensors are installed at the top of the link, at 90 degrees to each other, to measure the link\u2019s angular velocity. Figure 2 shows the details of the system components. The hardware components of the setup consist of motors, motor controllers, sensors, data acquisition cards, and processing unit. As motors, two flat brushless motors (EC 45, Maxon Motors) were chosen with nominal voltage 48 V, torque 134 mNm, and current 0.9 A. The diameter of each motor is 42.8 mm and its weight 140 g. To control the motors in current mode, a Maxon Digital EC controller (DEC) 50/5 was used. The gyro sensors on top of the link are Yaw Rate Gyroscope Sensor V1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003161_j.addma.2020.101345-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003161_j.addma.2020.101345-Figure8-1.png", + "caption": "Fig. 8. Schematics of the experimental setup.", + "texts": [ + " The 700 cross-sections include channels which have the same CAD dimensions and build direction but could not be cold flow tested. A representative microphotograph is depicted in Fig. 7. The measurements have been taken by visually marking the edges of the channel. Straight edges along the actual channel edges were approximated to measure the width and the height of each cross-section. In order to derive friction factors of the vast amount of channels, a cold flow test rig was developed. The setup is shown schematically in Fig. 8. Pressurized air is supplied by a compressor and a fluid conditioning unit ensures that the air is clean and dry. This is necessary in order to protect the mass flow meter from contamination as well as ensuring stable gas properties. The static pressure is adjusted and kept stable by a pressure control valve. The mass flow is then recorded by a mass flow meter, an Alicat M Laminar Flow Element from Alicat Scientific. The inlet temperature is measured using a thermocouple Type K. The differential pressure is measured in front of the sample using a Keller PRD-33 X differential pressure sensor", + " Different nozzle diameters are used for the appropriate channel sizes. The pressure tap is positioned 25dh downstream of the nozzle inlet to ensure the flow is fully developed to get a proper static pressure reading. Additionally, the circular channel in the nozzle extends 10dh after the pressure tap to ensure that the unsteady flow at the transition to the sample does not affect the differential pressure signal. The absolute ambient pressure is measured using a Keller PAA-33 X (not depicted in Fig. 8). The electrical signals (0\u202610 V, 4\u202620mA) from the sensors are read by an I/O module from Delphin Technology and transferred as digital signals via LAN cable. All sensors have been calibrated separately. Additionally, a commercially smooth pipe was used for reference measurements which agreed very well with the Colebrook correlation. To verify that the experiment is working correctly and consistently over all performed experiments, a reference experiment was performed before and after the experiments from which the reported data in this work was taken", + " The reference measurements which have been performed before and after the experiments are within a\u00b12%-band from the respective friction factor calculated from the Colebrook equation. As the measured mass flows are in a range of 0.05\u2026 0.40g/s even a very small leakage could impact the mass flow reading. Therefore, leakage checks are performed before and after each measurement session. For the leakage checks, the nozzle is covered by pushing it onto a flat surface. The system is filled with pressurized air and shut off using the stop valve (see Fig. 8). The differential pressure is recorded. A pressure loss of 5Pa/s is tolerated as this corresponds to a mass flow loss of < 1.5E\u22125g/s which is smaller than the resolution of the mass flow sensor and will therefore have no impact on the mass flow measurement. After the leakage check has been performed successfully the actual experiment can be started. The nozzle is centered on the channel to be measured. Then, the stop valve is opened. Proper sealing between the sample channel and the nozzle is verified using three separate methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003402_s11071-020-05852-8-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003402_s11071-020-05852-8-Figure6-1.png", + "caption": "Fig. 6 Beam NES free body diagram", + "texts": [ + " Several assumptions need to be made, namely that: (1) the beam is rigid, so all bending characteristics from its elastic physical features are assumed to be captured by the nonlinear spring characteristic; (2) the radius of curvature of the path described by the mass is constant, which is not true in reality as the length of the beam changes progressively as it wraps around the curved surface of the boundary; and, (3) the beam is massless, as its actual mass is much smaller than the concentrated mass at its tip. Consider the free body diagram shown in Fig. 6. The external forces that produce a moment around point B are the spring force ( fsp), damping force ( fd), and the gravitational force (mg), each one multiplied by the appropriate moment arm length to A. The vector equation of the acceleration of point B, accounting for its relative motion with respect to point A, is given by: aB = aA + \u03b1 \u00d7 rB/A \u2212 \u03c92rB/A = aA i\u0302 + \u03c6\u0308k\u0302 \u00d7 [\u2212l sin(\u03c6)i\u0302 + l cos\u03c6j\u0302] + [\u03c6\u03072l sin(\u03c6)i\u0302 \u2212 \u03c6\u03072l cos(\u03c6)j\u0302], (8) where rB/A is the position vector from point B to point A, \u03b1 is the angular acceleration vector of the beam, \u03c9 is the scalar angular velocity of the beam, and a is an acceleration vector of its corresponding point in the diagram indicated by the subscript", + " The base acceleration variable name is changed to u\u0308b, thus producing an equation in terms of angular moments, and assuming small angle approximations yields: (IB +ml2)\u03c6\u0308 + cl(l\u03c6\u0307) + fsp(l\u03c6) \u2212mgl(l\u03c6) = \u2212mlu\u0308b. (13) This constitutes the equation of motion (EOM) of this class of nonlinear spring in terms of angular states. For the present study, the gravitational effects included in Eq. (13) are neglected as the device is oriented in a horizontal plane, thus gravity points toward the paper plane in Fig. 6. To retrieve the motion states in the u direction, a simple substitution derived from the kinematics of the problem has to be made to the \u03c6 quantity, recalling that u = l\u03c6. Thus the angular and linear quantities are related by: u = l\u03c6, u\u0307 = l\u03c6\u0307, u\u0308 = l\u03c6\u0308. An approximation of the analytical solution of the developed EOMs is presented. A preliminary step in determining a set of adequate parameters for having a fully described system is also required. 3.1 Parametric design With the EOM and F-D characteristics completed, a parametric model is generated from physical constraints defined beforehand for this nonlinear spring, mainly ensuring that it would be easy to build, from commercially available materials, and that it shall produce relevant results when attached to existing physical components available in the laboratory, which are expected to be used during the experimental verification of the model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003047_icit45562.2020.9067209-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003047_icit45562.2020.9067209-Figure1-1.png", + "caption": "Fig. 1. Cross section of the analyzed IM.", + "texts": [ + " In this paper, also the cage reaction to the high order harmonics is included, considering both a skewed and unskewed motor. In particular, induced currents can be able to partially shield the flux penetration of the currents or slot harmonics, balancing the losses in iron and cage Joule losses. Depending on the skew angle the rotor response is more or less evident. The high frequency losses computation methods, presented in this work are based only on MS and Time-Harmonic (TH) FEA, in order to be computationally quick. II. Analyzed InductionMotor The lamination of the analyzed IM is reported in Fig. 1. The geometry is inspired by an automotive IM. In this example, the rotor cage material is aluminum and the cage bar geometry has been designed to get the lowest resistance with balanced magneto-motive force drop in the rotor with respect the stator. Even though the bar shape could be suitable for the skin effect, 978-1-7281-5754-2/20/$31.00 \u00a92020 IEEE 187 the rotor frequencies considered in a variable speed drive are not that high to affect the motor performance. Finally the cage end ring have been designed to have the lowest impact on the rotor resistance: the width takes all the available space below the stator end winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003796_icem49940.2020.9270813-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003796_icem49940.2020.9270813-Figure1-1.png", + "caption": "Fig. 1: Manufacturing process of the winding insertion and names of stator specimens.", + "texts": [ + " The windings are inserted manually into the stator cores. The winding is designed as a wound winding with w = 10 turns per coil and n = 24 number of parallel strands. The copper fill factor is kCu = 0.35. The winding is chorded with a coil pitch of y = 5. The end windings are pressed together to reduce the axial length of the machine and the distance of the end winding to the cooling sleeve added later. Important data like the length, the diameters of the stator cores are shown in table I. The type of stators and the winding process are shown in figure 1. The vibration behavior of a multi-mass system, such as the stator active part of an electric machine, can be described, using the differential equation M d2~\u03be(t) dt2 + D d~\u03be(t) dt + K~\u03be(t) = ~F (t), (1) where ~\u03be(t) is the deflection, M the mass matrix, D the damping matrix, K the stiffness matrix and ~F (t) the applied force [4], [18]. This system can be solved using the harmonic solution ~\u03be(t) = ~Xie \u03bbt. (2) The harmonic solution has i complex conjugated solution \u03bbi,1/2 = \u2212\u03b4i \u00b1 j2\u03c0fi. (3) that represent poles in the complex plane with the corresponding eigenmodes ~Xi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure18-1.png", + "caption": "FIGURE 18. EXPERIMENTAL SET UP FOR VBD BRUSH SEAL WEAR TESTING.", + "texts": [ + " The rig can be pressurized with either nitrogen (up to 0.35 MPa (50 psi)) or air (up to 3.45 MPa (500 psi)). Details of this test setup are described in [8] and [9]. For this test, sub-scale test seals were fabricated with a 127 mm (5\u201d) bristle pack diameter to fit the rotor size of the C5R. Cost-E and CrMo-V test rotors were machined and T-F-T straight cut, VBD brush seals were installed with a 0.254 mm (10 mil) interference with the test rotor. The rotors did not have any coating for this test. A cross section of the test setup is shown in Figure 18. In this setup, two VBD brush seals could be tested simultaneously by installing them on the drive end and the non-drive end of the rotor, respectively. For wear testing, the rotor is actuated at 1257 rad/s (12000 RPM) in ambient temperature at 2.07 MPa (300 psid.) This corresponds to a surface speed of 79.8 m/s (262 ft/s), which is above the operating speed of the seal in the end packing. High pressure room temperature air is introduced in the middle section and leaks out through the seals to downstream chambers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure3-1.png", + "caption": "Figure 3. Magnetic adhesion mechanism.", + "texts": [ + " One motor rotates the unit, and the other controls extension and magnet rotation. The servomotor for unit rotation is directly connected to the universal joint of the adjacent unit and controls the angle between the unit and its neighbor about the z-axis. Conversely, the servomotor for unit extension and magnet rotation is attached to the cam driver. The cam mechanism of the driver enables extension and contraction of the unit and magnet rotation. The structure of the magnetic adhesion mechanism installed in each unit is shown in Fig. 3. There is a permanent magnet in the center of the mechanism, as shown in Fig. 3 (b). This magnet can rotate inside the mechanism and become magnetized in its radial direction. In this way, the direction of its magnetic pole can be switched between the adhesion and roller side yoke. Magnet rotation allows the magnetic adhesion mechanism to switch between two magnetic circuit patterns, as shown in Fig. 4. In Pattern 1, the magnetic pole of the permanent magnet faces the adhesion yokes, which then establish a magnetic circuit through the adsorption surface. In this scenario, the magnet adhesion mechanism produces a large adsorption force. Further, the adhesion yokes are attached to friction plates (see Fig. 3 (a)). When the magnet adhesion mechanism adsorbs onto a surface, this structure imparts a high frictional force to the surface. In Pattern 2, the magnetic pole of the magnet faces the roller side yokes that form a magnetic circuit through the magnetic rollers and adsorption surface. The friction between the roller and the adsorption surface is small and allows the magnet adhesion mechanism to move easily. In addition, because the adhesion yokes are attracted to the magnet by leaked magnetic flux, which is generated by the gap between the adhesion and the rotation yokes, the adhesion mechanism separates from the adsorption surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000898_10402004.2013.830799-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000898_10402004.2013.830799-Figure6-1.png", + "caption": "Fig. 6\u2014Equivalent stiffness and damping of a lubricated contact.", + "texts": [ + " [18] The partial derivatives of xi and x\u0307i with respect to \u03c1i , xb, and yb are given by the following equations: \u2202xi \u2202\u03c1i = \u2202x\u0307i \u2202\u03c1\u0307i = \u03c1i \u2212 xb cos \u03b8i \u2212 yb sin \u03b8i xi [19] \u2202xi \u2202xb = \u2202x\u0307i \u2202x\u0307b = xb \u2212 \u03c1i cos \u03b8i xi [20] \u2202xi \u2202yb = \u2202x\u0307i \u2202y\u0307b = yb \u2212 \u03c1i sin \u03b8i xi . [21] In the EHD lubricated roller-to-race contacts, the damping is more effective in the inlet region due to the squeezing effect, whereas the stiffness is very high in the central zone due to the smaller film thickness. Therefore, the spring\u2013damper combination is considered to be in parallel connection (Sarangi, et al. (19)), as shown in Fig. 6. Under loaded condition, the total approach (\u03bb) of two contacting cylinders is given by \u03bb = \u03b4 \u2212 h, [22] where the elastic deformation (\u03b4) is calculated according to Harris\u2019s relation (Harris (21)) and h is the lubricant film thickness. Referring to Fig. 6, because the same load is carried by elastic deformation stiffness as well as by the lubricant stiffness and damping in parallel, the force balance can be written as F = KH\u03b41.11 + Ch\u0307 F = Kh + Ch\u0307 . [23] Further substituting h\u0307 = i\u03c9bh and using Eqs. [22] and [23], the total approach (\u03bb) can be expressed as \u03bb = ( F KH ( 1 \u2212 i\u03c9bC (K + i\u03c9bC) )) 1 1.11 \u2212 F (K + i\u03c9bC) , [24] where F is the contacting force, KH = 7.86\u00d7104\u00d7(L)8/9 N/mm1.11 is a constant of proportionality for load\u2013deformation relation, and K and C are the EHD stiffness and damping coefficients" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002186_elk-1410-48-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002186_elk-1410-48-Figure1-1.png", + "caption": "Figure 1. Speed-torque characteristics of parallel connected induction motor drives.", + "texts": [ + " Multiple induction motors controlled by a single inverter run in parallel in electric traction drives to reduce the cost, size, and need for maintenance. If the machines have the same speed-torque characteristics then speeds are equal, and torque-sharing rates are equal in all operating conditions. Practically, there will be observable differences between the behaviors of machines and the speeds may not be identical because of slight differences in wheel diameters. The speed-torque characteristics for slightly far from identical machines are shown in Figure 1 (dark blue \u2013 motor 2, green line \u2013 motor 1), which causes different torque sharing at the same speed. By assuming that the wheel diameter of machine 1 is a bit larger than that of machine 2, then the torque sharing of machine 1 will be higher in motoring mode, but lower in braking mode where the corresponding characteristic is represented in Figure 1 (blue line). Both the mismatch in characteristics and unequal wheel diameter problems exist in \u2217Correspondence: gunabalanr@yahoo.co.in real-time situations and the speed-torque behavior of both motors will differ and unbalance condition arises (red line \u2013 motor 1 and pink line \u2013 motor 2). If average currents flow through the stator windings and rotor fluxes are considered for unbalanced load conditions, the speeds of both motors deviate much from the command speed. To reduce the speed difference among the induction motors, average and differential currents were used to determine the reference currents [1,2] and the hardware results were presented only for step change in speed under no load conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure9-1.png", + "caption": "Figure 9. The relationship of curve-face gear and shaper cutter at point P.", + "texts": [ + "\u00f01\u00de !f !\u00f01\u00de \u00bcMf10 r \u00f010\u00de 1 \u00f0us, s\u00de !f !\u00f01\u00de \u00f017\u00de And the absolute velocity of non-circular gear can be represented as follows vf !\u00f02\u00de \u00bc rf !\u00f02\u00de !f !\u00f02\u00de \u00bcMf2 r \u00f02\u00de 2 \u00f0 k, 1\u00de !f !\u00f02\u00de \u00f018\u00de So put matrix (14) and equation (17), (18) into equation (12), the contact line L12 of non-circular gear can be represented as follows L \u00f0 f \u00de 12 \u00bc rss s \u00fe i1srss s\u00bdsin 2 \u00f0 os \u00fe s\u00de cos2\u00f0 os \u00fe s\u00de \u00fe i1srss sin\u00f02 os \u00fe 2 s\u00de \u00fe i1sL cos\u00bd 1 l\u00fe os \u00fe s \u00bc 0 \u00f019\u00de Derivation of the line contact of curve-face gear As shown in Figure 9, in the meshing process of curve-face gear and shaper cutter, the meshing equation (10) can be simplified as follows f \u00f0 2, 1, s\u00de \u00bc nf !\u00f01\u00de vf !\u00f013\u00de \u00bc 0 \u00f020\u00de Parameter Value Number of tooth of shaper cutter z1 12 Modulus of shaper cutter m (mm) 4 Pressure angle of shaper cutter 1 ( ) 20 Addendum coefficient of cutter gear ha * 1 Headspace coefficient of cutter gear C* 0.25 Eccentricity of non-circular gear e 0.1 Number of tooth of non-circular gear z2 18 Order of non-circular gear n2 2 Order of curve-face gear n3 4 Inner radius of curve-face gear R0(mm) 70 Pitch curve radius of curve-face gear R1(mm) 71", + "\u00f01\u00de is the unit normal vector and it is equal to nf !\u00f02\u00de, vf !\u00f013\u00de is the relative velocity of curve-face gear and shaper cutter in coordinate system Of XfYfZf. PC ! is the absolute velocity vector at point P of curve-face gear, and it is perpendicular to the radius vector O3P ! . PA ! is the absolute velocity vector at point P of shaper cutter, and it is perpendicular to the radius vector O1P ! . The relative velocity vf !\u00f013\u00de at point P can be represented by the two absolute velocity vectors shown in Figure 9. So, the relative velocity of point P can be established as follows v! \u00f012\u00de f \u00bc v! \u00f01\u00de f v! \u00f03\u00de f \u00bc PA ! PC ! \u00f021\u00de According to the unit normal vector nf !\u00f01\u00de of shaper cutter shown in Figure 3 and the transformation matrix Mf3 Table 3. Some contact points with no errors of alignment. 1 (rad) 0.0314 0.0628 0.0942 0.345 0.377 0.408 0.973 1.005 1.037 2 (rad) 0.0174 0.0347 0.0521 0.1894 0.2063 0.2231 0.5098 0.5248 0.5398 x 2.5917 2.5899 2.5862 2.4041 2.3564 2.2811 0.4584 0.5794 0.809 y 72.5011 72.5385 72" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002465_icsima.2014.7047443-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002465_icsima.2014.7047443-Figure1-1.png", + "caption": "Fig. 1. DOF helicopter manufactured by Quanser.", + "texts": [ + " Moreover, few researches focused on decentralized tracking control laws for MIMO models of uncertain plants [8]. The following part of this paper is organized into four sections. In Section II, we briefly describe the Quanser helicopter and its dynamics and mathematic model, followed by a formulation of the robust output regulation for the elevation and pitch axes of the 3-DOF helicopter. In Section III, a new design method of robust output regulator is proposed. Experimental results are presented in Section IV. A 3-DOF bench-top helicopter developed by Quanser shown in Fig. 1 and Fig. 2, where portrayed two DC motors has installed at the ends of a rectangular frame [9]. Those employed to control the 3-DOF lab helicopter. The helicopter frame can move freely to pitch around a long arm and the long arm suspended from a junction who has two degrees of freedom to elevate and travel. These two DC motors are called front motor and back motor separated with independent control signals. The helicopter frame would elevate if both motors were applied with positive voltages, while it would fall with opposite ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure14-1.png", + "caption": "Figure 14. Experimental setup to measure the dynamic frictional resistance of each gear type, including that with conical passive rollers", + "texts": [ + " The force gauge on the right showed that under every condition of pressing force, the t-test revealed significant differences between the static frictional resistance of the gear with conical passive rollers and other gear types. The advantage of this conical passive roller against static frictional resistance was thus confirmed in this experiment. In this research, the dynamic frictional resistances of the gear with conical passive rollers and other ordinary gears were also examined. The experimental setup used for this experiment is shown in Fig. 14. To achieve the constant sliding motion between each spur gear and the wider rack gear, we used the same normal rack and pinion gear mechanism shown in Fig. 12 and drove it with a geared AC motor, RSF-14B-100-F100-24B-C, manufactured by Harmonic Drive Systems Inc. Instead of the omnidirectional gear, we used a rack gear with wider teeth under the spur gear or the gear with passive rollers for the same reason as the static frictional resistance experiment described in the former chapter. We used the same spur and crowning gears as in the experiment to measure static frictional resistance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003042_10402004.2020.1737285-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003042_10402004.2020.1737285-Figure3-1.png", + "caption": "Figure 3. Test gear.", + "texts": [ + " Three new sets of AMS 6308 spur gears were run-in and tested under LoL conditions to different inflection points along the material\u2019s failure progression curve. The gears are fabricated from the same material batch, so it is assumed that they will behave in a similar manner when subjected to an LoL test. A thermocouple placed just above the gear mesh provided a close approximation of the temperature and health of the gears in real time. Posttest analysis included optical micrographs and surface profilometry of the tooth surfaces. Shown in Fig. 3 is a photograph of a test gear made from AMS 6308. AMS 6308 is a case-hardened steel that has high-temperature performance, exhibiting 61\u201363 HRC at 21 C and 55 HRC at 315 C. Its excellent toughness and fatigue resistance make it a popular choice for components in aerospace power transfer applications. The gears were manufactured to aerospace quality standards and were from the same material batch to avoid variations in batch-tobatch material properties that could influence the results. The nominal chemical composition of AMS 6308 is shown in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000855_kem.579-580.300-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000855_kem.579-580.300-Figure3-1.png", + "caption": "Fig. 3 The generated process of non-circular gear", + "texts": [ + " Order ))/(tan( BB xyabsa=\u03bb , so 1 \u03d5 can be obtained as follows, \u2265\u2264\u2212 \u2265\u2265+ \u2264\u2265\u2212 \u2264\u2264 = )0,0(,2 )0,0(, )0,0(, )0,0(, 1 BB BB BB BB yx yx yx yx \u03bb\u03c0 \u03bb\u03c0 \u03bb\u03c0 \u03bb \u03d5 . (6) The arc length that cylindrical gear roll from A to B can be calculated as follows. \u222b += g AB dddrrS 0 22 )/)(()( \u03d5\u03d5\u03d5\u03d5 . (7) Thus, the angle caused by the pure roll is rS AB / 2 =\u03d5 , and then the total turned angle \u03d5 of cylindrical gear is 21 \u03d5\u03d5\u03d5 += . The coordinates of 1 o point in )-( xyoS can be calculated as follows[7], \u2212= \u2212= B B yggry xggrx o o )sin()( )cos()( 1 As shown in Fig. 3(a), gear blank is fixed and the initial position of center point of cylindrical gear is coincident with the origin of polar coordinate of non-circular gear. At the every discrete polar angle of non-circular pitch curve, the cylindrical gear turn the angle \u03d5 and move to the point ),( 11 OO yx . Divided equally the pitch curve as equal polar angle or equal arc length, and then let gear billet and cylindrical gear make solid geometry operation at every discrete polar angle. The generating cutting process when cylindrical gear roll purely along the pitch curve is also the simulation analysis of gear shaping, the generated four-order elliptic gear is shown in Fig. 3(b). The Distribution Control of Gear Teeth. During the transmission of non-circular gear pair, the teeth of driving gear and the teeth space of driven gear have one-to-one relationship, so it\u2019s necessary to set the teeth distribution based on the initial assembly position of gear pair. Calculating the polar radius maximum max1 r of driving gear and the polar radius minimum min2 r of driven gear, and then the assembly center distance of non-circular gear pair is min2max1 rrd += . Order driving gear has a tooth at the polar angle 1g at which the polar radius is 1r , and then the polar radius of driven gear corresponding to driving gear is 12 rdr \u2212= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure6-1.png", + "caption": "FIG. 6. Plan view schematic of a disk with radius R and a small hole of radius r located at the center. The three radial creases are shown dashed. Each face between two creases is subdivided into 2n + 1 facets which each subtend an angle \u03b1. Mirror symmetry is assumed across the face bisector and only one half of a face, shown shaded, is analyzed.", + "texts": [ + " While this precludes the localized dimple observed in experiments, it corresponds to the assumptions of previous studies on the inversion of f-cones [14,15]. We therefore find the consequences of this assumption on the predicted mechanical behavior. For conciseness the analysis is restricted to symmetric behavior, such that all creases deform simultaneously and identically. However, the analysis can be adapted for general deformations. Consider an initially planar thin disk with thickness t , radius R, and a small hole of radius r located at the center, as shown in Fig. 6. N equally spaced radial creases are imposed dividing the disk into N identical faces. Therefore, only one face, of angle 2\u03c0/N , is considered for this analysis. This face is divided by 2n equally spaced radial hinges, forming 2n + 1 wedge-shaped facets, each subtending an angle of \u03b1 = 2\u03c0/N (2n + 1). Mirror symmetry is assumed across the face bisector and only one half is analyzed, shown shaded in Fig. 6. The facets are all joined at the central vertex but are otherwise 043001-3 unconstrained. We number the facets clockwise, with facet i located between hinges i \u2212 1 and i. The centerline of facet n + 1 is aligned with the line of symmetry. A Cartesian coordinate system is located at the center of the undeformed disk with the y axis aligned radially along a crease and the z axis pointing out of the page. Initially, the facets on each side of a crease are rotated about the crease line by an angle of \u03b2/2 from the x-y plane", + " Bending crease If bending is allowed along the crease line, then an alternative deformation mode characterized by a localized dimple around the indentation point occurs, as described in Sec. II. Following the rigid-faceted approach, we discretize each face 043001-7 between creases into a series of rigid facets and postulate a kinematic mechanism which captures the dimple deformation. The postulated facet layout is shown in Fig. 13. There are two key differences in comparison to the straight-crease model (Fig. 6). First, there is a circumferential hinge located at a radius of , measured along the crease from the center of 043001-8 the disk. Second, there are additional facets which emanate from the intersection of this hinge with the crease line (facets C and D). We assume that the facets directly adjacent to the crease (A\u2013F) have subtend angles of \u03b10, while all other facets have subtended angles of \u03b1 = (\u03c0 \u2212 2\u03b10)/(2n \u2212 1). Like the straight-crease model, we consider the kinematics before deriving the strain energy of deformation for this model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003673_tmag.2020.3035280-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003673_tmag.2020.3035280-Figure18-1.png", + "caption": "Fig. 18. 12s/10r DA-FSPM prototype machine. (a) Stator and rotor laminations. (b) Machine structure.", + "texts": [ + " 17(a), if the stator conductors are pushed from the slot opening to 60% of the slot depth, the torque deviations of the SLMC and FLMC models are increased while that of the TLMC model is reduced. The accuracy of the FLMC model is improved to some extent when the rotor conductors are pushed from the slot opening to 67% of the slot depth, as shown in Fig. 17(b). However, the accuracy of the SLMC and TLMC models is reduced. Nevertheless, in general, the TLMC model still exhibits the highest accuracy. A 12s/10r DA-FSPM machine is manufactured and tested to validate the LMC models and FEM analysis. Table V and Fig. 18 show the key parameters and structure of the prototype machine. It should be noted that the stator iron segments are linked by 0.5 mm iron bridges in order to ease the fabrication, whereas they are neglected in the LMC model. As shown in Fig. 19, the torque predicted by 2-D FEM is slightly smaller than that predicted by the LMC models due to increased PM leakage flux caused by the iron bridges. The average torque predicted by 2-D FEM is larger than the 3-D FEM predicted Authorized licensed use limited to: San Francisco State Univ" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001730_978-3-319-18296-4_13-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001730_978-3-319-18296-4_13-Figure2-1.png", + "caption": "Fig. 2 Explanation for the calculus of infinitesimal element of the wheel contour", + "texts": [ + " The slope of the tangent M\u03c4 is given as tanu \u00bc dy dx : \u00f04\u00de By differentiating relations (2) and (3), it is obtained dx \u00bc \u00f0q cos h\u00fe q0 sin h\u00dedh \u00f05\u00de dy \u00bc \u00f0q sin h q0 cos h\u00dedh; \u00f06\u00de where q0 \u00bc dq dh : \u00f07\u00de Inserting the above differentials in (4), it reads dy dx \u00bc tanu \u00bc q sin h q0 cos h q cos h\u00fe q0 sin h : \u00f08\u00de The geometrical considerations drive to the angle between the OwM and the normal M\u03bd b \u00bc h u: \u00f09\u00de When the wheel rolls without slipping on the rail, the initial contact point O between the wheel and the rail arrives in O\u2032, the point M comes in the point M\u2032 on the rail head; when the wheel center Ow arrives in the point Ow\u2032 (Fig. 1b), the abscissa of the M\u2032 point is xr. We are interested in calculating the coordinate (xw, yw) of the wheel center when the wheel turns around the center point at the angle \u03b8. To this end, the length of the arch OM has to be known. According to the geometrical correlation in Fig. 2, the length of an infinitesimal element of the wheel contour is ds \u00bc q cos b dh \u00f010\u00de and the length of the wheel contour between O and M points is s \u00bc Z ds \u00bc Zh 0 q cos b dh \u00f011\u00de When the wheel rolls on a rigid rail and the wheel rotation angle is \u03b8, the distance between the initial contact point O and the current wheel\u2013rail contact point M\u2032 and the arch length are equal xr \u00bc s: \u00f012\u00de According to Fig. 1b, the coordinates of the wheel center are as follows xw \u00bc xr q sin b \u00f013\u00de yw \u00bc q cos b: \u00f014\u00de In the preceding geometrical considerations, we tacitly suppose that the wheel and the rail have a single contact point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002574_detc2016-60019-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002574_detc2016-60019-Figure10-1.png", + "caption": "Fig. 10 CAD model and a 3D layered based virtual AM model", + "texts": [], + "surrounding_texts": [ + "For the Data transfer in coupled operations, MATLAB 14 programming language is used as a source system to write the contour lists file. The contour list data with correct data structure should be constructed for the successful coupling along with the CAD environment. Here, the target system is SOLIDWORKS macros. Therefore, contour list data will be written in a macros files format (.SWP files) as required in SOLIDWORKS. The process of writing of .SWP file data structure for the contour points and the corresponding height of extrusion or layer thickness includes following steps. Firstly, the contour lists are generated. Then, the vertices of each contour and corresponding constant extrusion height are written according to the .SWP file structure as shown in Fig. 7. Figure 7 displays the data format of .SWP file based on sliced contours lists. Set Part syntax opens a new part document in SOLIDWORKS software; Set skSegment syntax is used for modelling the contours and Set myFeature syntax is used for extrusion-purpose. e.g., if a first 2D contour has n points (i.e., P1, P2, P3, P4,\u2026Pi\u2026,Pn) the second contour has m points (i.e., P1, P2, P3, P4,\u2026Pj\u2026,Pm) at two different heights then according to .SWP file structure these should be written as Set skSegment = Part.SketchManager.CreateLine(P1, P2, P3, P4,\u2026Pi\u2026,Pn) Set skSegment = Part.SketchManager.CreateLine(P1, P2, P3, P4,\u2026Pj\u2026,Pm) Lastly, the complete file structure of 3D layered model from 2D contour lists in the form of .SWP file format is generated through MATLAB program and is used in the data transfer in coupled operations from the source system (i.e. MATLAB R2014a) to target system (i.e. SOLIDWORKS 2012)." + ] + }, + { + "image_filename": "designv11_34_0002902_j.cirpj.2020.01.001-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002902_j.cirpj.2020.01.001-Figure3-1.png", + "caption": "Fig. 3. Graphical representation of the problem being solved.", + "texts": [ + " The dependence of total emissivity on the steel surface temperature was used in the model for a wide range of temperature changes in accordance with the published data [22]. A heat transfer condition by heat conduction into the tooling was proposed for the bottom surface. If the node indexed (x,y,z\u20131) gets into the tooling, we can use the expression lf ix Ttx;y;z Tchamber Sf ix \u00bc l\u00f0Tt 1=3 x;y;z \u00de T t x;y;z Ttx;y;z\u00fe1 Dx \u00f07\u00de where lfix = 400 W/(m K) is the thermal conductivity of a copper fixture and Sfix is the fixture thickness - assumed to be 0.05 m. A graphical representation of the problem being solved is shown in Fig. 3. The average diameter of the layer deposited on a massive substrate is 72 mm (the outer diameter of the cylinder is 78 mm and the inner diameter is 66 mm). The height of one layer being deposited is 1 mm, and the width is 6 mm. Since during the experiments the electron beam was moving in a circle with the frequency of 100 Hz, the heating source was represented as a oints: (a) the start of the process; (b) during the first layer deposition; (c) during the g zones in additive manufacturing by means of electron beam metal 0", + " Wire feed was simulated by adding control volumes having an environmental temperature of 25 C in front of the heating source. In the model, the cross section of the added material was assumed to be equal to the cross section of the layer being deposited. Deposition speed was 340 mm/min. Using Eqs. (1)\u2013(7), a computer program was created in the Microsoft Visual Studio environment. The time step was set to 0.05 s and the step of the spatial grid was set to 0.5 mm. During all computational experiments, temperatures were recorded at points having indices \u2013100 . . . 10, as shown in Fig. 3. All points were located in the center of the beam\u2019s motion trajectory. The length of the bead between the points 100 and 0 is 1/4 of the length of the whole circle layer. Point 0 coincides with the point around which the beam oscillates. Point 10 is located at approximately 5,6 mm behind the beam and point 10 is located at 5,6 mm in front of the beam. Between the points 10 and 10 were also selected control points with the indices 8, 6, 4, 2, 0, 2, 4, 6, 8 (they are not shown in the Fig. 3) equidistant from each other. Please cite this article in press as: D.A. Gaponova, et al., Effect of reheatin wire deposition method, NULL (2020), https://doi.org/10.1016/j.cirpj.20 Fig. 4 shows the mathematical modeling results in a graphical form. It can be seen that a stationary regime is not observed because of the deterioration of heat removal from the product and change in massiveness of the body. The molten pool length\u2014represented by grayscale\u2014increases from 10.5 mm when the first layer is deposited, to 18 mm when the third is complete. Fig. 5, a shows the temperature-time dependencies for each of the points (T\u2013100, T\u201350, T\u201310, T-2, T0, T5, T10) in Fig. 3\u2014which were stored at each time step. Beam power during the computational experiment was constant and assumed to be 1.5 kW. It can be seen from the above dependences that molten pool temperature increases significantly during the deposition process. As such, it is necessary to introduce temperature feedback. Temperature fluctuations in the calculated dependences, especially noticeable for high values (for example, at the point T-2), are associated with the adopted spatial discretisation. (The cylindrical structure in the model consists of cubic volumes)", + " The difference of these signals is fed into the controller of the beam current regulator, which acts on the bias supply source connected to the electron gun control electrode. The proposed method was developed into a mathematical model. As a temperature regulator, the simplest proportional regulator was chosen\u2014determined according to the formula where K is the proportionality constant of the regulator. The value of constant correction, C, was selected experimentally. As the temperature Tmes had to be stabilized, the temperature T\u201310 at the point of index \u201310 was chosen (Fig. 3). The set value, Tset, was chosen as 1400 C. Additionally, temperatures were recorded for points 1 mm below those shown in Fig. 3 (to understand their location, the figure shows a point with index -100). Specified temperatures are dashed and indexed as T\u2019\u2013100 . . . T\u201910. Temperature variation over time (when temperature feedback is introduced) is shown in Fig. 9, a. This data were obtained using the model with K = 0.001 and C = 450. To reduce C, a switch from proportional to a proportional-integral control law is possible, but these studies have not yet been carried out. In the model of closed loop system the stabilization starts at t = 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001802_978-3-662-46312-3_29-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001802_978-3-662-46312-3_29-Figure3-1.png", + "caption": "Fig. 3 Fine blanked helical gear and detailed view of the 100 % smooth sheared tooth flank", + "texts": [ + " The calculation of the cutting force is modified as shown in Eq. 2: Fs D s cos \u02c7 lsRmcs: (2) The determined experimental data of the cutting force correspond very well with the calculated cutting force and confirms the mathematical approach of Eq. 2. The feasibility study of the fine blanking of helical gears was carried out with a casehardened steel 16MnCr5 and a sheet thickness of 3.2 mm. In Table 1 the helical gear is specified in detail. The fine blanked helical gears showed a very good quality, see Fig. 3. However, the quality of the tooth flank differs for the right and left flanks. The right flank shows a smooth sheared edge of 100 % and the left flank of only 81 %. Furthermore, the analysis of the die roll height gives different results along the tooth profile. For the left flank a value of 0.12 mm was determined and in comparison to the right flank (0.65 mm) the difference is nearly by factor 5. These results indicated a change of the die clearance during the fine blanking process [7]. The change of the die clearance is undesired and therefore the experimental work was supported by the application of the finite element method to determine the cause-and-effect relationships between the part quality and fine blanking parameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001946_ecce.2014.6953971-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001946_ecce.2014.6953971-Figure1-1.png", + "caption": "Fig. 1 IPMSM geometry, ref. HW820CG.", + "texts": [ + " So iron losses may be comparable to the copper losses and could have a real impact on the machine performances and the sizing machine. Consequently, in an optimization process of sizing machine, a fast and accurate iron losses model is required. In [1], Internal Permanent Magnet Synchronous Machines (IPMSM) appear to be the most suitable actuators for these applications because of their high power density and their high efficiency. This machine type will be studied in this paper with the geometry is given in Fig. 1. Firstly, this paper will introduce the state of art of models calculating the iron losses in materials. As the field weakening influences the Flux Densities Waveforms (FDW) especially in the teeth and in the yoke, the proposed model calculates the FDW from a semi-analytical approach, i.e. a nonlinear nodal network with magnetic saturation [11]. The phenomena linked to the speed rotor and the control laws, i.e. FDW, are decoupled in a polynomial formulation of the 978-1-4799-5776-7/14/$31.00 \u00a92014 IEEE 4188 iron losses in function of the fundamental frequency and non-constant coefficient depending of the control laws", + " In field weakening operation, we cannot assume a sinusoidal waveform of the flux density so we will use models taking into account the influence of the flux density derivative on iron losses as LS and generalized Bertotti models. Although the LS model gives a good description of the magnetic losses, the generalized Bertotti model could be more suitable because it is faster and easier to calibrate with the material type. The developed model for our application obtains the FDW from a nonlinear nodal network with the methodology described in [11] and solving the problem in magneto-static. The IPMSM of Fig. 1 can be discretized in different areas as described on Fig. 3. These areas represent the main flux ways as described in Fig. 2 which have been defined from a preliminary study of the IPMSM with finite elements software (FLUX 2d). The nodal network has been defined from a first study with FE approach is given in Fig. 4. So this model cannot provide the temporal FDW in the different parts of the stator but the spatial FDW. However it is possible to assume that the temporal FDW in the teeth and the stack are identical with the spatial FDW" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002189_s10544-016-0092-9-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002189_s10544-016-0092-9-Figure11-1.png", + "caption": "Fig. 11 Schematic showing a top-view and b side-view of fabricated FET devices with silicone well suitable for analyte testing", + "texts": [ + " The corresponding relative standard deviation also decreased from 80.5 % to 16.5 %. The decrease in electrical resistivity could be due to the increased percolation pathways between flakes from the new growths. The ethanol CVD treatment decreased the standard deviation of inter-measurement sites significantly, thus solving a critical problem for the practical use of RGO in electronics such as the development of biosensors (Faulkner et al. 2014; Huang et al. 2013b; Huang et al. 2013c) and microelectronics (Su et al. 2009; Huang et al. 2014b). Figure 11 shows a schematic of a field-effect-transistor (FET) biological sensor platform that could be fabricated on top of the post-treated GO transducer reported in earlier sections. Gold electrodes are deposited on the transducer. Then, a silicone rubber well is fixed on the electrodes. The liquid analyte of interest can then be placed into the silicone well and interact with the transducer. Different liquid analyte will interact differently with the transducer and result in different FET behaviour. This behaviour can be recorded in a semiconductor parametric analyser" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002247_jsee.2016.00070-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002247_jsee.2016.00070-Figure2-1.png", + "caption": "Fig. 2 Missile/target engagement geometry", + "texts": [ + " Here, we consider the impact angle, denoted by imp\u03b8 , is the angle be- tween velocity vectors of the missile and the target as in [21]. The impact angle geometry can be seen in Fig. 1. Denoting the heading angles of the missile and the target at the terminal time as \u03b8Mf and \u03b8Tf, respectively, the impact angle, \u03b8imp, is defined as \u03b8imp= \u03b8T f \u2212\u03b8Mf. (1) Here, for the case of stationary targets, we define the heading angle of the target as \u03b8Tf= 0. Consider a planar engagement between the missile and the slowly moving target as shown in Fig. 2, where (xM, yM) and (xT, yT) denote the position of the missile and the target, respectively. The missile heading angle, velocity, and lateral acceleration are expressed by \u03b8M, VM and AM, respectively. Similarly, the target heading angle, velocity, and lateral acceleration are expressed by \u03b8T, VT and AT, respectively. The lateral acceleration command is normal to the velocity vector of the missile. Since only non-maneuvering targets are considered, the target lateral acceleration AT= 0. q is the LOS angle and R is the relative distance between the missile and the target", + " The first simulation is the SMC guidance law with various guidance gains in order to analyze the guidance gains affecting the guidance performance. The second simulation analyzes the FSMC guidance law for the lag-free system using the on-line optimized guidance gains technique. The third simulation involves testing the FSMC guidance law for the first-order lag system. Finally, nonlinear simulations of the FSMC guidance law are compared with other guidance laws with terminal impact angle and acceleration constraints. The homing geometry as described in Fig. 2 with following parameters is used to setup the test engagement scenario. Firstly, we outline parameters for homing geometry. The initial position of the missile is (0, 10 000) m, and the velocity is 300 m/s. A stationary target is placed on (15 000, 0) m. The SMC guidance law parameters are defined as follows: k1=2, k2=2, and \u0394 =0.01. The lateral acceleration command bound is chosen as AMmax=80 m/s2. The defined error distance is assumed to be df = 3 m. To demonstrate the influences of guidance gains, five different guidance gains\u2014SMC-10, SMC-30, SMC-50, SMC-70, SMC-90\u2014are used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001068_s1068375513030022-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001068_s1068375513030022-Figure2-1.png", + "caption": "Fig. 2. Specimens for the coating. On the right\u2014metal block blank for specimen, on the left\u2014specimen ready for ESA.", + "texts": [ + " In addition to the indicated mechanical parameters (the specimen feed rate, amplitude and frequency of the transverse oscillations of the vibrator), the ALIER 31 installation itself made it possible to vary the mode of operation, the energy coefficient and amplitude of the electrode vibration. Thus the experimental facility provided an oppor tunity to vary modes by 6 parameters during the mech anized (hence, with stabilized parameters) coating process. SURFACE ENGINEERING AND APPLIED ELECTROCHEMISTRY Vol. 49 No. 3 2013 ELECTROSPARK ALLOYING FOR DEPOSITION ON ALUMINUM SURFACE 183 The methods for direct coating deposition were as follows. For the specimens, alloy D1 (GOST 4784) was used as the substrate. The specimens in a rectan gular form (Fig. 2) were cut from D1 alloy sheets for friction and wear tests. After the milling cut the specimens surfaces were polished with an abrasive cloth strips to diminish roughness before the ESA coating. After polishing and marking the specimens were weighed on a VLR 200 analytic balance, and afterwards they were fixed with two screws on the sliding carriage of the experimental facility. The electrode from the Al\u2013Sn alloy (~20 mass % of Sn, manufactured by the above method) was also preliminary weighted and fastened in the vibration generator of ALIER 31", + " Method 2 presupposes the usage of a con stant value of the \u201csupplied charge\u201d or an amount of energy introduced into the discharge gap during the deposition process (in the present study it is refered to as the method \u201cwith constant energy amount\u201d). Using method 2, the number of the plated coatings was increased proportionally to an increase in the TE feed rate. According to method 1, the constant number of the deposited layers was 4 at various TE feed rates with respect to the specimen; the rates varied in the range of 0.3\u20132 mm/s: the TE traveled as many as four times with a preset rate along the entire surface under treat ment (Fig. 2). In the case of deposition by method 2 four layers were deposited (the TE traveled 4 times along the surface) at the speed of 0.5 mm/s of the TE dis placement and 16 layers at the feed rate of 2 mm/s; 8 and 12 layers were deposited at the feed rates of 1 and 1.5 mm/s, respectively. The results of the weighting were used to determine a specific deposition rate G in mg/(s cm2) which, depending on conditions, could be either of a positive or of a negative value; in the latter case, the weight of the specimen after the treatment did not increase but rather decreased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001564_j.optlastec.2014.05.017-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001564_j.optlastec.2014.05.017-Figure2-1.png", + "caption": "Fig. 2. PUCU (a) top view; (b) 3D sombrero-profile; and (c) geometrical characteristic of cross-section along line in (a).", + "texts": [ + " Especially, pulse laser is widely used because it can produce higher peak power, which is not only beneficial to break through the damage threshold of high reflective materials but also helpful to reduce the heat affected zone of strengthened layer [17]. When single pulse laser beam radiates onto the material, molten pool is formed instantly. Under the synthetic action of phase transformation, surface tension and/or vaporization [16,18,19], a unit possessing different surface morphology, microstructure or chemical composition from the original surface is produced in the remelting area. This unit, which is similar to the construction unit on the body surface of natural organisms and embedded into the substrate like a \u2018peg\u2019, as shown in Fig. 2, is named as peg-shaped bionic coupling unit (PBCU). In industrial applications, the quality of PBCU can be defined in terms of geometrical characteristics, which consist of structural and morphological factors. The structural features of PBCU can be described using the dimensions of diameter (D), penetration depth (PD), cross sectional area (CSA), and the surface morphological characteristics mainly involve the surface profile with peak and valley, that is maximal bulge height (BH) and innermost cup depth Contents lists available at ScienceDirect journal homepage: www", + " In the experiment, medium carbon steel was chosen as the experimental matrix and its chemical composition is analyzed using the instrument of Energy Dispersive Spectrometer (EDS), as given in Table 3. Specimens of 35 25 6mm3 were cut by electric spark machine. In order to avoid the influence of surface roughness on the absorptivity of laser, the surface of specimens were mechanically polished to remove the machine marks and make metal surface at the same roughness (Ra 0.847 \u03bcm), then degreased in acetone. PBCU with morphological and structural features, as shown in Fig. 2, was fabricated by single pulse gauss laser beam from a 800W (rated power) Nd:YAG pulsed laser machine as per experimental arrangement listed in Table 2. Fig. 3 shows the schematic diagram of process and detailed process will be described in the Section 5. To measure the geometrical characteristic, each PBCU was wire-cut along the diameter. Metallographic samples were prepared in accordance with standard procedures used for metallographic preparation of metal samples. This involved coarse polish using 300 , 600 , and 1000-grit silicon carbide (SiC) impregnated emery paper followed by fine polishing using 5 \u03bcm and 1", + " In turn, a downward flow is generated in the center of the molten pool [32]. Within the scope of experimental parameters, the evolution process of BH and CD are piecewise with the augment of processing parameters. When the processing parameters are moderate, the circular flow in molten pool is unaffected or slightly affected by the vaporization of metal, thus the movement of material is mainly dominated by the surface tension. As a result, liquid metal is pulled toward the center of the molten pool, PBCU with sombrero-profile, as shown in Fig. 2, is formed. As is well known to us, power density and pulse duration have significant effect on the temperature of molten pool [31]. Combining the formula (20), it is easy to understand that the decrease of DA and/or increase of PP and/or PD can elevate the temperature of molten pool. According to the relationship between the surface tension gradient and the temperature of molten pool [35], \u2202s \u2202y \u00bc cp 1\u00fe ln T T0 \u2202T \u2202y \u00f023\u00de where cp represents the specific heat, T0 the reference temperature used for calculating surface enthalpy, T the temperature of molten pool, s the surface tension and y the vertical coordinate in the cross-section of molten pool, higher temperature of molten pool can lead to larger surface tension gradient, as a result, more liquid metal is pulled toward the center of the molten pool, which causes the initial increase of BH and CD" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002946_j.precisioneng.2020.03.001-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002946_j.precisioneng.2020.03.001-Figure15-1.png", + "caption": "Fig. 15. Schematic of the top specimens and the base block: (a) specimen, (b) base block.", + "texts": [ + " In this section, we report a measurable effect of the distribution of the real contact areas on contact stiffness and that this was verified by experiment. An experiment was conducted to measure changes in the contact stiffness. In this experiment, contact surfaces with different distributions of the real contact areas were made using the CMC method. The contact stiffness of each contact surface was measured for comparison. The specimens were pushed onto a base block to make a contact surface. Fig. 15 shows a schematic of the top specimens and the base block. The specimens had different distributions of the real contact areas because the specimens had different amplitudes of waviness. Fig. 16 shows the experimental setup for measuring contact stiffness. The horizontal and vertical contact stiffness was measured for each specimen. A vertical preload was applied using a bolt and measured using force sensor 1. Horizontal and vertical forces were applied to the specimen using a piezoelectric element" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003704_ecce44975.2020.9236376-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003704_ecce44975.2020.9236376-Figure8-1.png", + "caption": "Fig. 8 Three-phase series-end winding topology for torque mode", + "texts": [ + " Drive the corresponding thyristors of the topology for speed mode or torque mode. Change to the corresponding modulation method. (a) 2022 Authorized licensed use limited to: Linkoping University Library. Downloaded on June 19,2021 at 13:14:34 UTC from IEEE Xplore. Restrictions apply. V. SIMULATION AND EXPERIMENT The proposed topology for torque mode and the structureswitched method have been verified by simulation and experiments. The proposed drive system with induction motor (IM) are established as shown in Fig. 8. The parameters of the IM are listed in Table I. For the proposed drive, the switching frequency is set to 20kHz and the leg dead time is set to 1\u03bcs. The DC bus voltage is set to 100V. TABLE I MACHINE PARAMETERS OF THREE-PHASE IM Machine parameters Value Number of stator pole pairs(P) 1 Excitation Inductance (Lm) 0.2692 H leakage inductance (Lm) 9.5mH Stator resistance (Rs) 2.876\u2126 rotor resistance (Rs) 1.274\u2126 Rated rotation speed (N) 1000RPM Rated phase current (IN) 10 A Rated Power (Prate) 2.2kw Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001923_insi.2012.55.8.417-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001923_insi.2012.55.8.417-Figure1-1.png", + "caption": "Figure 1. Modelling of gear tooth crack: (a) modelling of cracked tooth; (b) tooth notation; (c) uniform crack distribution", + "texts": [ + " Different crack propagation scenarios were studied in[11], where it was found that the crack model of a constant crack depth along the whole tooth width resulted in the largest effect on the RMS and kurtosis, used as indicators. Accordingly, to study how the FRFs change with mesh stiffness reduction due to crack propagation, it is more significant to consider the crack propagation scenario, which involves the highest reduction in the mesh stiffness. Therefore, in the current study, the calculation model presented in[5] was applied under the assumption that the crack extended through the whole tooth width with a constant crack depth, as explained in Figure 1. Gear tooth crack detection using dynamic response analysis Omar D Mohammed and Matti Rantatalo are with the Division of Operation, Maintenance and Acoustics, Lule\u00e5 University of Technology, Sweden. Tel: +46 7253 90777; Email: omar.mohammed@ltu.se Omar D Mohammed is also with the Mechanical Engineering Department, College of Engineering, Mosul University, Iraq. Email: omar.mohammed@ uomosul.edu.iq Paper presented at CM 2013 & MFPT 2013, The Tenth International Conference on Condition Monitoring and Machinery Failure Prevention Technologies, Krak\u00f3w, Poland, June 2013 DOI: 10.1784/insi.2012.55.8.417 By considering the tooth as a non-uniform cantilever beam with an effective length of \u2018d\u2019, see Figure 1(b), the deflections under the action of the force can be determined and then the stiffness can be calculated. Note that, in this part of the stiffness calculations, the crack is assumed to have a constant crack depth, q0, along the tooth width, W, as explained in Figures 1(a) and 1(c). Based on a calculation of the potential energy stored in a meshing gear tooth, it is feasible to obtain the bending, shear and axial compressive stiffnesses as follows[5]: 1 Kb = xcos \u03b11( )\u2212 hsin \u03b11( )( )2 E Ix0 d \u222b dx ....................(1) 1 Ks = 1.2cos2 \u03b11( ) G Ax0 d \u222b dx ............................(2) 1 Ka = sin2 \u03b11( ) E Ax0 d \u222b dx ...............................(3) where: Kb is the bending stiffness, Ks is the shear stiffness and Ka is the axial compressive stiffness. h, hc, hx, x, dx, d and \u03b11 are illustrated in Figure 1(b) (\u03b11 varies with the gear tooth position). E is the Young\u2019s modulus and \u03c5 is the Poisson ratio. G is the shear modulus, G = E 2 1+\u03c5( ) . Ix is the area moment of inertia, Ix = 1/12( ) hx + hx( )3 .W , hx \u2264 hq 1/12( ) hx + hq( )3 .W , hx > hq \u23a7 \u23a8 \u23aa \u23a9 \u23aa . Ax is the area of the section of distance \u2018x\u2019 measured from the load application point, Ax = hx + hx( ).W , hx \u2264 hq hx + hq( ).W , hx > hq \u23a7 \u23a8 \u23aa \u23a9\u23aa . hq = hc \u2013 q(z) sin(\u03b1c); hq represents the reduced dimension of the tooth thickness when a crack exists. q0 and \u03b1c are the crack depth and crack angle, respectively, as shown in Figure 1(b). We can find the tooth stiffness resulting from the effect of all the stiffnesses calculated previously as follows: Ktp = 1 1 Kb + 1 Ks + 1 Ka \u239b \u239d\u239c \u239e \u23a0\u239f ...........................(4) stiffness Sainsot et al[12] studied the effect of fillet foundation deflection on the gear mesh stiffness, derived this deflection and applied it for a gear body. The fillet foundation deflection can be calculated as follows[12]: \u03b4 f = F.cos2 \u03b1m( ) W .E L\u2217 u f S f \u239b \u239d \u239c \u239e \u23a0 \u239f 2 + M \u2217 u f S f \u239b \u239d \u239c \u239e \u23a0 \u239f + P \u2217 1+Q\u2217 tan2 \u03b1m( )( ) \u23a7 \u23a8 \u23aa \u23a9\u23aa \u23ab \u23ac \u23aa \u23ad\u23aa " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003261_tii.2020.3004389-Figure21-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003261_tii.2020.3004389-Figure21-1.png", + "caption": "Fig. 21 Thermocouples for hotspot detection in the windings.", + "texts": [ + " The PMSM is mounted on a large power motor test bench and connected to a dynamometer gantry load. As shown in Fig. 20, The PMSM is under control of an integrated inverter controller and cooled by the water flow with a constant temperature. The three-phase input current is recorded by high-precision current transducers. To obtain the temperature of hotspot in the PMSM and testify the prediction accuracy of the proposed model, two sets of Kthermocouples are embedded at the end and active winding parts of the PMSM. Fig. 21 presents the installation position of the thermocouples in the windings. Besides, to verify the local accuracy of the model, a sufficient number of thermocouples are placed on the iron core, housing, endcap, end cavity, and bearing of the motor, as shown in Fig. 22. Most thermocouples are inserted by drilling small holes and buried by thermal grease. The readings of the thermocouples are recorded by a data acquisition board (Smacq-PS2016). To implement this algorithm for thermal management, the proposed theoretical model is first constructed and debugged in MATLAB" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003065_s40684-020-00217-3-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003065_s40684-020-00217-3-Figure6-1.png", + "caption": "Fig. 6 Tool paths for a roughing and b finishing", + "texts": [ + " PowerMill is a professional CAM software package for complex high-speed and five-axis machining. Before the tool path was generated, the work range of the five-axis hybrid machine tool with the laser module was established to prevent interference by the laser module and the measuring device during the machining process. With regard to the NC code generation result, the roughing time was four and a half hours and finishing time was two hours (the total machining time was six and a half hours). Figure\u00a06 shows the tool paths for the roughing and finishing steps. Figure\u00a07 presents the experimental set-up, including the measuring device and production process of the turbine blade. The five-axis hybrid machine tool used in this study was a Hyundai WIA Hi-V560M vertical M/C with a laser module. A Laserline high-power diode laser (HPDL, LDM1000-100) with a wavelength range of 940\u2013980 nm was used. The working distance (focal length) was 138 mm (length between the laser and specimen). It is important to maintain the working distance because the temperature can decreased by more than 100\u00a0\u2103 when error of \u00b1 1 mm in the working distance is generated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002946_j.precisioneng.2020.03.001-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002946_j.precisioneng.2020.03.001-Figure9-1.png", + "caption": "Fig. 9. Setup of the experiment.", + "texts": [ + " Kono Precision Engineering 63 (2020) 197\u2013205 observation of the plastic deformation was conducted at high magnification using a laser microscope. Table 2 shows the specifications of the specimens. The specimens were made of carbon steel (JIS S45C) and the surfaces were plated with Y. Jorobata and D. Kono Precision Engineering 63 (2020) 197\u2013205 nickel phosphorus (Ni\u2013P). The specimens were machined using shaper cutting with a single crystal diamond tool. Specimen 1 was machined with waviness about 1 \u03bcm in amplitude, and specimen 2 was machined with a 5 \u03bcm step. Fig. 9 shows the setup of the experiment. The top specimens were pushed onto the base specimen using the CMC method. The 350 N preload was applied to the contact surfaces with bolts. The preload was measured with a force sensor. The specification of the force sensor is shown in Table 1. After unloading the preload, the contact surface shape was measured using a laser microscope. Fig. 10 shows the allowable waviness under the experimental conditions (W \u00bc 350 N, h \u00bc 10 \u03bcm, R \u00bc 0.2 mm pm \u00bc 540 kgf/mm2). The calculated allowable waviness was 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure9-1.png", + "caption": "Fig. 9 Machining tolerance", + "texts": [ + " (14) and (21) should be expressed with r and \u03b8 in polar coordinate system. The Y coordinate of IEi under polar coordinate system is, y \u00bc r \u22c5 cos \u03b8 So, when yA 1, we have inequalities \u2212 \u03c0 < \u0394\u03d5 < 0. (27) Hence, the body rotates about axis Cx3 in the direction opposite to the direction of rotation of vector \u03c1 which is chosen positive in Fig. 2. It follows from inequalities (27) that the magnitude of angle \u0394\u03d5 does not exceed \u03c0 . However, if we need to turn the body by an angle greater than \u03c0 , we can rotate it by a smaller angle in the opposite direction. Vector \u03c1 can perform several revolutions along the circle of radius a. If N is the number of revolutions, then the magnitude of the angle of rotation for body P + Qo is given by formula |\u0394\u03d5| = N\u03c0(D \u2212 1) D (28) which follows from (26). Let us substitute formula (25) forD into Eq. (28) and solve the obtained equationwith respect to radius a" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003753_jestpe.2020.3037942-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003753_jestpe.2020.3037942-Figure11-1.png", + "caption": "Fig. 11. Open-circuit magnetic field distributions. (a) Conventional SFPM. (b) C-core SFPM. (c) E-core SFPM. (d) CP-SFPM.", + "texts": [ + " Downloaded on June 03,2021 at 19:52:45 UTC from IEEE Xplore. Restrictions apply. 2168-6777 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. respectively. The torque contributions of the working harmonics in the four SFPM machines, are summarized and tabulated in Tables II~V. The open-circuit magnetic field distributions of the four SFPM machines are shown in Fig. 11. Compared with the conventional, C-core and E-core structures, the CP-SFPM machine exhibits significantly reduced flux leakage in its stator yoke, which is mainly due to the fact that the additional iron bridge located on the outer stator yoke provides an effective magnetic circuit for multiple field harmonics. It can be found that the PMs play a part of flux barriers for the |3ZPM-Zr| (5th) order harmonic in the conventional SFPM machine, as shown in Fig. 11(a), resulting in relatively lower low-order working harmonic amplitudes and more seriously flux leakages in stator yoke. Similarly, the flux barrier effect can also be observed for the |5ZPM-Zr| (2nd) order harmonic in the C-core SFPM machine, as illustrated in Fig. 11(b). Compared with the C-core structure, the E-core design exerts a relatively lower magnetic reluctance for the |5ZPM-Zr| (2nd) order harmonic, as shown in Fig. 11(c). As a result, the E-core machine has higher |5ZPM-Zr| (2nd) order harmonic amplitude and the corresponding torque contribution, as tabulated in Table IV. In addition, the CP-SFPM machine with iron bridge design, as reported in [24] [25], provides an additional magnetic circuit for low-order working harmonics so that the lower-order working harmonic amplitude, i.e., |2ZPM-Zr| (1st), and torque capability can be significantly improved compared with the other SFPM machines. Some key findings can be summarized as: 1) Due to the prominent amplification effect of the gear ratio Grj, the dominant torques are contributed by the low-order air-gap field working harmonics in the four SFPM machines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003631_lra.2020.3033258-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003631_lra.2020.3033258-Figure2-1.png", + "caption": "Fig. 2. Three-link robot model.", + "texts": [ + " The latter, which will be the focus of this study, is a more complicated yet more general approach. In recent years, crawling is exploited extensively [11] in the rapidly growing bio-inspired field of soft robotics [12]. In these robots, which draw inspiration from nonskeletal animals, actuation is often similarly created by pressurizing fluids in internal cavities of their compliant structure. Nonetheless, rapid dynamic movement of soft robots can also be achieved \u2013 with other actuation methods, such as magnetic fields [13]. In previous works [14] we have shown that a three-link model (Fig. 2) captures well the major phenomena of quasistatic inchworm crawling of a soft bipedal robot. This lumped model is a very basic form of multi-contact bipedal crawling, in analogy to McGeer\u2019s biped [15] being a benchmark of bipedal walking. Yet, most \u201ctraditional\u201d articulated legged robots are of higher complexity and often have more legs [16], [17]. To the best of our knowledge, the minimal three-link robot has not been noticeably studied. The purpose of this study is to model and analyze the dynamic bipedal frictional inchworm crawling locomotion, in order to comprehend and exploit the effects of inertia, for both soft and articulated robots", + " We also investigate the effects of friction uncertainties, which was shown to have major influence [20], and mass asymmetry, propose a novel input shaping technique and apply machine-learning based optimization to improve the performancein traveling distance. Finally, we discuss a feedback control strategy. In order to investigate crawling at frequency range where inertial effects are significant, we have manufactured the three-link robot prototype in Fig. 1. The robot has a central link with length l0, mass m0 and moment of inertia J0 and two distal links, with length li, mass mi and moment of inertia Ji, for i = 1, 2 (see Fig. 2). For most of this work we consider identical distal links l1 = l2 \u2261 l, m1 = m2 \u2261 m, J1 = J2 \u2261 J , and in Section VII we investigate the influence of asymmetric mass distribution. The experimental setup parameters are summarized in Table I. Two servomotors at the joints receive a sequence of angle commands from the microcontroller (in open-loop) and track it with internal closed-loop control. Assuming planar motion and point-contacts, a corresponding three-link model is proposed in Fig. 2. The motion of the robot can be described by the generalized coordinates q(t) = [x y \u03b8 \u03d51 \u03d52] T, where (x, y, \u03b8) are the planar position and absolute orientation angle of the central link, and (\u03d51, \u03d52) are the joint angles. Throughout this work we assume that the two joint angles qc(t) = [\u03d51(t) \u03d51(t)] T are prescribed directly as known periodic input functions (and hence their time-derivatives q\u0307c(t), q\u0308c(t) are also known). This assumption, which is quite reasonable for a robot with joints controlled in closed-loop, significantly simplifies the analysis by effectively reducing the number of dynamically-evolving degrees-of-freedom (DoF)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure3-1.png", + "caption": "Figure 3: Staggered bristle layout for an unloaded seal", + "texts": [ + " The bristles were made of Haynes 25, which is typically preferred in most of turbine sealing applications due to their superior strength and satisfactory ductility up to 600oC. Cold worked Haynes 25 material with 10% cold reduction has 725 MPa tensile yield strength and 1070 MPa ultimate tensile strength limits at room temperature [10]. Seal photos are given in Figure 1, and brush seal dimensions are detailed in Figure 2. Using the axial spacing of 25% and tangential spacing value of 5%, the number of bristle rows in seal axial direction is found to be 15 by using the formula for a staggered bristle layout given by Aksit [7] (also visualized in Figure 3). 2 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use A free-state stiffness measurement system has been designed to characterize the seal stiffness under unpressurized static rotor conditions. The measurement relies on pressing a metallic pad, which has an arc surface of the same diameter as the seal inner diameter, against unloaded bristles. As detailed in Figure 4, the metallic pad movement is generated by the linear slide" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000408_ccece.2015.7129153-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000408_ccece.2015.7129153-Figure1-1.png", + "caption": "Fig. 1. Helmholtz coils . A Helmholtz coil is a device for producing a region of nearly uniform magnetic field. A Helmholtz pair consists of two identical circular magnetic coils that are placed symmetrically along a common axis, one on each side of the experimental area, and separated by a distance h equal to the radius R of the coil. Each coil carries an equal electrical current flowing in the same direction.", + "texts": [ + " Compared with the permanent magnet, the electromagnetic coils are more flexible to control and alter the magnetic field. Therefore electromagnetic coils are widely used for actuation. There are two types actuation force that can be used for remote manipulation, the magnetic torque and the magnetic force. Magnetic torque makes the magnet rotating and the magnetic force pull or push the magnet. There are two kinds of coils based actuation methods, 1) the uniform magnetic field and constant gradient magnetic field, which are generated by Helmholtz coils (as shown in Fig.1) and Maxwell coils (as shown in Fig.2), respectively [6]\u2013[9]. The Helmholtz and Maxwell coils usually work together to control a micro magnetic particle. The model of these two coils are very simple and therefore the system is easy to be realized. The uniform field from Helmholtz coils provide the torque and the constant gradient from Maxwell \u2217Corresponding authors are Ren H. and Yu H., mail:{ren, bieyhy}@nus.edu.sg This work is supported by Singapore Academic Research Fund, under Grant R397000139133, R397000157112, R397000156112 and NUS Grant C397000039001" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000759_cjme.2014.0827.142-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000759_cjme.2014.0827.142-Figure2-1.png", + "caption": "Fig. 2. Arc arrangement and calculation domains of TWIA", + "texts": [ + " Energy conservation equation: CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7125\u00b7 2 lam tur( ) ( )\u2022 ,\u2022 R p k kh J vh h S t C \u00e6 \u00f6\u00f7\u00e7 \u00f7\u00e7 \u00f7\u00e7 \u00f7\u00e7 \u00f7 +\u00b6 + = \u00f8 - + \u00b6 \u00e8 (3) where h is the enthalpy, \u03c3 is the electrical conductivity, SR is the net radiative emission coefficient, lam turk k+ is the thermal conductivity and turbulence effects included, and Cp is the specific heat. Current continuity equation: 0.J\u00b4 = (4) Ohm\u2019s Law: .J =- (5) Maxwell\u2019s equation: 0 .B J\u00b4 = (6) Lorentz force: .F J B= \u00b4 (7) Current density: ( ).J E v B= + \u00b4 (8) Electric field: . A E t \u00b6 =- - \u00b6 (9) Where J is the current density, \u03c6 is the electrical potential, \u03bc0 is the permeability of vacuum, v is the plasma velocity, and A is the magnetic field strength. TWIA is symmetric about the plane which formed by the axis of the twin wires. Thus, a half model was established, as shown in Fig. 2. Due to the larger change of arc parameters near the polar region, a non-uniform grid-point was employed. Finer meshes were employed in polar region and a relatively coarse grid in arc column region. Symmetry boundary conditions were exerted on ABCD. The boundary condition of current density on cathode surface (MN) is as follows: 2 ,C I J R = (10) where R is the radius of the cathode surface. The axisymmetric arc model has been widely studied, and the critical boundary condition is the current density at the cathode tip: some, such as HSU, et al[24], prefer to input a current density profile, while others choose to include the cathode in the calculation domain and set current density at the top cylindrical part of cathode[25\u201326]", + " The final calculation result, 20000K, was obtained after several iterations and not influenced by the initial assumed temperature. At present the cathode region and anode region cannot be described accurately by the modeling method and two element(1mm thick) results was employed to describe the arc parameter in cathode region and anode region. TWIA was divided into two parts: the polar region and arc column region. The polar region can be further divided into two parts: cathode region(1 mm thick, area KTWL in Fig. 2(b)), anode region(1 mm thick, area HEFI in Fig. 2(b)). The arc column region can also be divided into two parts: that between the cathode region and anode region in the horizontal direction(upper part of arc column region) and the remaining part(lower part of arc column region). The nodal solutions on the model\u2019s plane of symmetry were subsequently extracted. Such nodal position details were described by SHI, et al [29]. The TWIA calculated plasma temperature distribution on a symmetry plane is shown in Fig. 4. It can be seen that the highest plasma temperature of 17758K occurs close to the anode tip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003756_1350650120966894-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003756_1350650120966894-Figure7-1.png", + "caption": "Figure 7. Contact ellipse under load.", + "texts": [ + " Calculation of parameters relevant to sliding friction coefficient on tooth surface These parameters mainly include length of contact line, load per unit contact line, comprehensive curvature radius for contact position, relative sliding velocity and entrainment velocity for contact position. Calculation of length of contact line Under load, contact point becomes contact ellipse. The major axis of ellipse is much larger than the minor axis of ellipse and then it is still taken as con- tact line, which is described in Figure 7. In the process of solution, the contact line is discretized into n points (Figure 5). Coordinate of these discretized points can be obtained and then the length of contact line can be calculated. Calculation of load per unit contact line The contact line (major axis of ellipse) is discretized into n points (Figure 5). By the solution of LTCA, the load on discretized point (Figure 8) can be obtained, respectively. The load total of all discretized point constitutes the load on contact line (equation (1))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001925_ceidp.2014.6995760-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001925_ceidp.2014.6995760-Figure2-1.png", + "caption": "Fig. 2. Detailed model of composite croSS-aim", + "texts": [ + " The length of phase A is 10 m with an installation height of 38 m; the length of phase B is 11 m with an installation height of 33 m; the length of phase C is 10 m with an installation height of 28 m. The middle part of the tower is mainly made up of four vertically arranged post insulation rods. These post insulation rods are divided into four sections by three intermediate flanges. Adjacent post insulation rods are connected by oblique insulation rods. The structure of composite cross-arms in three phases is similar. As shown in fig. 2, the main part of the composite cross-arms is a two V -shaped post insulators, and two upward sloping V -shaped tension insulators, Post insulators and tension insulators are connected by connecting fittings and connected to the tower by connecting fittings. Post insulators are connected with conductors by yoke plate. The conductor used here is four bundle, the subconductor diameter is 26.8mm and the bundle diameter is 400 mm. Considering the interaction between three phases, the harmonic analysis method was used to obtain the maximum electric field strength in one entire power frequency cycle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001426_s0219876215500061-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001426_s0219876215500061-Figure1-1.png", + "caption": "Fig. 1. Solidworks model of gear\u2013pinion set.", + "texts": [ + " Based on the numerical results, the polynomial functions, the polynomial response surface (PRS) is utilized to formulate the optimal objective \u03c3max. Then, the polynomial function model of \u03c3max is used as the surrogate models in the stress-relieving holes optimization. Finally, the genetic algorithm (GA) methodology is used to find out the optimal locations and sizes of the holes where the maximum stress at the root fillet is reduced to the maximum limit. A numerical example illustrates the feasibility of the optimization design. Consider a gear-pinion set whose geometric model is shown in Fig. 1. In order to develop the internal stress-relieving method by design of stress-relieving holes, structural analysis based on the finite element (FE) method is used to calculate the 1550006-3 In t. J. C om pu t. M et ho ds 2 01 5. 12 . D ow nl oa de d fr om w w w .w or ld sc ie nt if ic .c om by N A N Y A N G T E C H N O L O G IC A L U N IV E R SI T Y o n 04 /2 7/ 15 . F or p er so na l u se o nl y. 2nd Reading November 7, 2014 16:31 WSPC/0219-8762 196-IJCM 1550006 P. Van Thoan et al. structure behaves of the gear under certain load configurations [Andrews (1991)]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002592_kem.725.378-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002592_kem.725.378-Figure1-1.png", + "caption": "Fig. 1 The pattern diagram of the movement of the nozzle during printing the specimen", + "texts": [ + " (#72752399, Queen's University, Kingston, Canada-06/01/17,06:51:59) thicknes of the layer of 0.2 mm, the nozzle temperature of 458 K, the bed temperature of 353 K, 363 K and 373 K, and the printing rate of 120, 420 and 1080 mm/min. The printing conditions of the ABS specimen were the thicknes of the layer of 0.2 mm, the nozzle temperature of 518 K, the bed temperature of 363 K, and the printing rate of 1000 mm/min. The pattern diagram of the movement of the nozzle during printing the specimen is shown in Fig. 1. Fig. 1 (a) shows the first layer and Fig.1 (b) shows the seconde layer. The nozzle movement pattern is as follows, as the first layer, (1) the nozzle drow 3 laps on the perimeter and holes of the specimen. (2) fill the inside with 45\u00b0 angle respect to the direction of tension force. As the second layer, repeat (1), then fill the inside with -45\u00b0 angle respect to the direction of tension force which is cross direction of (2). After that, first layer and second layer are repeated till the objective height. In this paper, the thickness of the specimen is 1 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000099_2017030-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000099_2017030-Figure3-1.png", + "caption": "Figure 3. Second model: Slanted spring.", + "texts": [ + " In this way, for a compressed spring, we recover a sort of Coulomb law, since the friction coefficients are proportional to the normal force exerted by the spring on the surface, that, in the limit, is exactly equal to \u03b1. Moreover, if the profile of the fluctuations is asymmetric, in the sense that \u03c9+ = \u03c9\u2212, then also the friction is asymmetric. Our second model generalizes the first one, since in this case we consider a slanted spring forming a fixed angle 0 < \u03d1 < \u03c0/2 with the vertical axis, as illustrated in Figure 3. As before, one end of the spring has fixed height and horizontal position u(t). In this case, however, the horizontal position of the second end will be different from u(t) and denoted with p(t). We therefore have u \u2212 p h \u2212 w\u03b5(p) = tan \u03d1. (3.7) We can express explicitly u as a function of p as u \u2212 h tan\u03d1 = p \u2212 w\u03b5(p) tan \u03d1. (3.8) We require \u03c9+ < cot\u03d1, (3.9) so that w\u2032 \u03b5(p) tan \u03d1 < \u03c9+ tan \u03d1 < 1 and therefore p\u03b5(u) is a one-to-one correspondence. The length of the spring is thus L = \u221a (u \u2212 p)2 + (h \u2212 w\u03b5(p))2 = u \u2212 p sin \u03d1 = h \u2212 w\u03b5(p) cos\u03d1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002484_12.2243652-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002484_12.2243652-Figure1-1.png", + "caption": "Figure 1. Setup and relationship of coordinates in robot polishing system.", + "texts": [ + " In this paper, we introduce the setup of the robot polishing system in our lab, and present the results of material remove experiments in grinding and polishing. The progress of two freeform surface fabrication is shown in section 4, one is a toroid concave surface and the other is a off-axis parabolic surface. The grinding and polishing results are presented and discussed here. The whole robot polishing system consisted of a 6-axises industrial robot, a turntable for holding workpiece, and the polishing tool, as shown in figure 1. A TX200 model industrial robot from St\u00e4ubli is used in our system. During polishing, the workpiece is fixed on the turntable, and the tool is carried by robot arm to move across the whole AOMATT 2016: Advanced Optical Manufacturing Technologies, edited by Wenhan Jiang, Li Yang, Oltmann Riemer, Shengyi Li, Yongjian Wan, Proc. of SPIE Vol. 9683, 96832D \u00a9 2016 SPIE \u00b7 CCC code: 0277-786X/16/$18 \u00b7 doi: 10.1117/12.2243652 Proc. of SPIE Vol. 9683 96832D-1 Downloaded From: http://proceedings.spiedigitallibrary", + " This method cannot be used in deterministic optical polishing because there is huge number of points in the moving path. So the off-line programming method is used in robot optical polishing. The VAL3 language is used to control the TX200 robot. So we wrote software in VAL3 which runs on the robot controller in order to interpret the G-code like polishing path file to the robot moving code. The polishing path file is generated by our own polishing software. The coordinate systems of the robot polishing setup are show in figure 1. The industrial robot has a base coordinate system named as world coordinates which located at the root of robot, and a default tool coordinate system named as flange coordinates which located at the end of robot arm. The frame coordinates is used to describe the position and posture of the workpiece in the world coordinate system. On the other hand, the tool coordinates is used to describe the position and posture of the polishing tool head relative to the flange coordinate system. The specifications of TX200 are list below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000914_20130825-4-us-2038.00108-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000914_20130825-4-us-2038.00108-Figure5-1.png", + "caption": "Figure 5. Digging kinematics defined by bucket angle steering; and joystick configuration example for the bucket steering kinematics, (upper left insert).", + "texts": [ + " The movement of the bucket of the Bobcat 435 prototype in Cartesian kinematics is shown in Figure 3. The joystick signals were emulated by the software for repeating straight line motions along three overlapping squares in rotated angles by 45 o Figure 4 shows the recorded trajectories of the bucket edge for the free-hand control of the excavator without Cartesian coordinate system rotation. On the basis of the Cartesian kinematics, a new digging and loading kinematics is introduced as bucket steering by Danko (2007). The concept is illustrated in Figure 5. Figure 2. Block diagram for kinematics transformation. Figure 3. Bucket movement control in Cartesian virtual kinematics. Coordinate system is at 45\u00b0 slope angle. Figure 4. Bucket movement in Cartesian virtual kinematics by free-hand operation. The trajectory is defined as the path of the bucket edge while the bucket moves forward in the tangential, xdirection of a Cartesian coordinate system of variable slope angle. The angle of the x-coordinate to the horizontal plane is controlled by the operator as the direction of the bucket movement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003444_j.matpr.2020.07.218-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003444_j.matpr.2020.07.218-Figure3-1.png", + "caption": "Fig. 3. Arc constriction mechanism [16].", + "texts": [ + " It is recognized that halides helps in electrical interactions which influence the physics of arc where as oxides controls the movement of metal flow in weldments (Marangoni convection movements). Fluorides affect the arc constriction mechanism because the affinity of fluorides with electrons is higher which directly relates the capacity of flux to be combining with electrons and influence arc constriction mechanism [25,32,33]. The Schematic of effect of arc constriction mechanism is shown in Fig. 3 [16]. The A-TIG performance is dependent on various process parameters such as Flux Type, Forces acting during the process, parameters used for process and type of material used as work piece [34]. The activated flux TIG welding process utilized for various types of ferrous and non ferrous metals [17,35,36]. The effect of Cu2O, NiO, Cr2O3, SiO2 and TiO2 were investigated on SUS 304 using GTAW process. The oxygen content increases the depth to width proportion if it is in the range of 70\u2013300 ppm, also change the surface tension gradient which results in increased penetration [37]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000560_cca.2015.7320681-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000560_cca.2015.7320681-Figure1-1.png", + "caption": "Fig. 1. Quadrotor-pendulum Model", + "texts": [ + " Another contribution is the easiness of the controller design since we apply only offset-type input transformation and no coordinate transformation differently from [13]. Finally, the validity of the present law is demonstrated via simulations, where we explicitly show the effectiveness of the vertical input by comparing with the LQR control case. II. 3D QUADROTOR-PENDULUM MODEL We first build the dynamic model of a 3D inverted pendulum on a quadrotor based on the Euler-Lagrange 978-1-4799-7787-1/15/$31.00 \u00a92015 IEEE 513 equation [19]. Throughout this paper, we consider the 3D quadrotor-pendulum model shown in Fig. 1. The position and orientation (roll, pitch and yaw angles) of the quadrotor frame \u03a3V in the world frame \u03a3O are respectively represented by [x y z]T \u2208 R 3 and (\u03b3, \u03b2, \u03b1), \u03b3, \u03b2, \u03b1 \u2208 (\u2212\u03c0, \u03c0]. On the other hand, the orientation of the pendulum on the quadrotor is given by (\u03c6, \u03b8), where \u03c6, \u03b8 \u2208 (\u2212\u03c0, \u03c0] represent the rotation around the x and y-axes on the world frame \u03a3O, respectively. We now derive the motion equation of the quadrotorpendulum system. We first suppose that the quadrotor can directly input the vertical body acceleration a \u2208 R and angular body velocity [\u03c9x \u03c9y \u03c9z] T \u2208 R 3, which might be implemented by a local controller. This assumption is imposed by aiming at the application to actual machines and also used in many other works on quadrotors. In Fig. 1, the parameters mp, Lp, L (= Lp/2), g > 0 mean the mass and length of the pendulum, the length from the pivot point to the center of mass of the pendulum and the gravity acceleration, respectively. Suppose also that any friction is small enough to be negligible. Then, the Euler-Lagrange equation [19] gives the following 3D quadrotor-pendulum dynamic model. \u23a1 \u23a3x\u0308y\u0308 z\u0308 \u23a4 \u23a6 = O V R(\u03b3, \u03b2, \u03b1) \u23a1 \u23a300 a \u23a4 \u23a6+ \u23a1 \u23a3 0 0 \u2212g \u23a4 \u23a6 , (1) \u03c6\u0308 = 3 2Lp cos \u03b8 ( y\u0308 cos\u03c6+ (z\u0308 + g) sin\u03c6 ) + 2\u03b8\u0307\u03c6\u0307 tan \u03b8, (2) \u03b8\u0308 = 3 2Lp (\u2212x\u0308 cos \u03b8 \u2212 y\u0308 sin\u03c6 sin \u03b8 + (z\u0308 + g) cos\u03c6 sin \u03b8 ) \u2212 \u03c6\u03072 sin \u03b8 cos \u03b8, (3)\u23a1 \u23a3\u03b3\u0307\u03b2\u0307 \u03b1\u0307 \u23a4 \u23a6 = \u23a1 \u23a3 \u03c9x cos\u03b1\u2212\u03c9y sin\u03b1 cos \u03b2 \u03c9x sin\u03b1+ \u03c9y cos\u03b1 \u2212\u03c9x cos\u03b1 tan\u03b2 + \u03c9y sin\u03b1 tan\u03b2 + \u03c9z \u23a4 \u23a6 ", + " Therefore, by minimizing \u03c8, we can hope for stronger control inputs and, as a result, better attenuation of \u2016q\u2016 than LQR control for large \u2016q\u2016. Also, because the weight for \u03bd(q) is the inverse of the weight for the vertical input uz (i.e. r\u22121 z ), the controller applies larger uz to achieve smaller \u03bd(q). This is consistent with intuition and implies the efficiency of the vertical input uz . The image of the cost function \u03c8 is illustrated in Fig. 2. IV. VERIFICATION We finally demonstrate the present control scheme (7) and (12) via 3D simulations. The parameters of the quadrotorpendulum system in Fig. 1 are set as mp = 1.0 [kg], Lp = 1.0 [m] and g = 9.8 [m/s2]. Then, we apply the present control law (7) and (12) to the quadrotor-pendulum model (5) from the initial pendulum orientations \u03c6(0) = \u22120.15 [rad] and \u03b8(0) = 0.1 [rad] (all the other initial states are set as 0). Here, we set the state weight matrix Q as the identity matrix and the input weight matrix as R = diag{1, 1, 0.01}. Moreover, to make the vertical input uz (i.e. a\u0303) effect obvious, we especially set much smaller rz than the other weights in this simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003443_j.procir.2020.04.140-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003443_j.procir.2020.04.140-Figure8-1.png", + "caption": "Fig. 8. Clearance vector loop model of an additively manufactured, non-assembly 8-bar mechanism", + "texts": [ + " The adaption of the procedure for solving the vector closure equations from Fig 4 for solving the equations for the statistical tolerance analysis using sampling techniques is shown in Fig. 7. As a result, the mobility band for the 4-bar mechanism can be calculated whereby the functionality can be ensured, evaluating the previously defined FKC. 4. Case Study The previously presented methodology is now applied to a case study of an additively non-assembly manufactured mechanism, consisting of eight linkages and seven clearance affected joints (see Fig. 8). By applying equation (1) to the case study, it becomes evident that two vector loops are now required for solving the vector loop model: L1 \u00b7 ei\u00b7\u03b81 + c12 \u00b7 ei\u00b7\u03b312 + L2a \u00b7 ei\u00b7\u03b82+ +c23 \u00b7 ei\u00b7\u03b323 \u2212 LAB \u00b7 ei\u00b7\u03b8A \u2212 L3 \u00b7 ei\u00b7\u03b83 = 0, (7) L1 \u00b7 e j\u00b7\u03b81 + c12 \u00b7 e j\u00b7\u03b312 + e j\u00b7\u03b82 (L2a + L2b \u00b7 e\u2212i\u00b7\u03b82a2b )+ +c23 \u00b7 e j\u00b7\u03b323 + c24 \u00b7 e j\u00b7\u03b324 + L4 \u00b7 e j\u00b7\u03b84 + c45 \u00b7 e j\u00b7\u03b345+ L5 \u00b7 e j\u00b7\u03b85 \u2212 LAC \u00b7 e j\u00b70 = 0. (8) In order to determine the clearance vector for the enhancement of the vector loop model, as described previously, a MBS is neglected", + " The adaption of the procedure for solving the vector closure equations from Fig 4 for solving the equations for the statistical tolerance analysis using sampling techniques is shown in Fig. 7. Fig. 7. Procedure for the statistical tolerance analysis using sampling technique As a result, the mobility band for the 4-bar mechanism can be calculated whereby the functionality can be ensured, evaluating the previously defined FKC. 4. Case Study The previously presented methodology is now applied to a case study of an additively non-assembly manufactured mechanism, consisting of eight linkages and seven clearance affected joints (see Fig. 8). By applying equation (1) to the case study, it becomes evident that two vector loops are now required for solving the vector loop model: L1 \u00b7 ei\u00b7\u03b81 + c12 \u00b7 ei\u00b7\u03b312 + L2a \u00b7 ei\u00b7\u03b82+ +c23 \u00b7 ei\u00b7\u03b323 \u2212 LAB \u00b7 ei\u00b7\u03b8A \u2212 L3 \u00b7 ei\u00b7\u03b83 = 0, (7) L1 \u00b7 e j\u00b7\u03b81 + c12 \u00b7 e j\u00b7\u03b312 + e j\u00b7\u03b82 (L2a + L2b \u00b7 e\u2212i\u00b7\u03b82a2b )+ +c23 \u00b7 e j\u00b7\u03b323 + c24 \u00b7 e j\u00b7\u03b324 + L4 \u00b7 e j\u00b7\u03b84 + c45 \u00b7 e j\u00b7\u03b345+ L5 \u00b7 e j\u00b7\u03b85 \u2212 LAC \u00b7 e j\u00b70 = 0. (8) In order to determine the clearance vector for the enhancement of the vector loop model, as described previously, a MBS P" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001297_j.jmatprotec.2013.01.009-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001297_j.jmatprotec.2013.01.009-Figure4-1.png", + "caption": "Fig. 4. Example of thermal field", + "texts": [ + " However, the simulation method cannot model these parameters I and h directly. To be representative, the heat source parameters P and R must be linked to the welding parameters I and h. In this study, two methods identifying the relation between the process parameters (I,h) and the heat source input parameters (P,R) are presented. 2.3.3. Mesh Thanks to conditions of symmetry, only a half plate was modeled. The number of elements was about 30,000. The element size was 0.2 mm in the finest zone, 0.6 mm in the surrounding zone and 1 mm elsewhere (Fig. 4). 2.4. Procedure The aim of this study was to define a numerical heat source which can predict the thermal field in the part. It was chosen to correlate the experimental (I,h) and numerical (P,R) data by comparing use of identical numerical and experimental torch speed. To quantify discrepancies between the experimental and the numerical results, the melted zone dimensions were used; a distinction was made between non-penetrating and penetrating welds: obtained during welding. on-pe - - 3 n f ( 3 p a l t s w b c p w m 3 s l l f l o e h e 4 w e L and l) of the experimental test produced the values of (P,R) corresponding to the experimental weld bead dimensions Lexp, and eexp (or Lexp and lexp)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003524_j.addma.2020.101612-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003524_j.addma.2020.101612-Figure14-1.png", + "caption": "Fig. 14. Roundness of the produced rods after (a) SLM process and (b) laser rounding process.", + "texts": [ + " An uneven roundness with a rougher side view can be observed in the figure. Fig. 13(b) shows the cross section of the rounded rod product. The experimental conditions for the rounding process were identical to those used in the SLM process. The roundness and longitudinal roughness have been significantly improved. The average roughness, Ra, of the rod produced by SLM decreased from 15 to 8.7 \u03bcm after the laser rounding. The roundness of the rod after the rounding process was compared using an image-analysis system. Fig. 14(a) shows the roundness of the product obtained from the SLM process, whereas Fig. 14(b) shows that obtained from the rounding process. The rounding process improves the rough surface contours that are visible after SLM and enhances the roundness of the rod. The rounding process improves the rough surface contours that are visible after SLM and enhances the roundness of the rod. The circularity tolerance zone size of the rounded rod decreases from 82 \u339b to 66 \u339b and the surface quality is improved. Fig. 13(c) shows the cross section of the sized rod product. After sizing the rounded rod through a 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001286_cdc.2014.7040359-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001286_cdc.2014.7040359-Figure1-1.png", + "caption": "Fig. 1. Agents\u2019 positions and orientation angles", + "texts": [ + " Section V derives stability conditions and illustrates them with some examples. Section VI concludes the paper. We consider a collection of n agents aj , j = 1, . . . , n in the plane, arranged cyclically so that aj+n is identified with aj . We denote by lj the distance between the agents aj\u22121 and aj ; by \u03c6j the oriented angle from the horizontal axis (with positive direction) to the ray \u2212\u2212\u2212\u2212\u2192ajaj+1; by \u03b1j the oriented angle from the ray \u2212\u2212\u2212\u2212\u2192ajaj+1 to the velocity vector of aj ; and by \u03c8j the angle aj\u22121ajaj+1 (see Figure 1). We note that the values of the angles \u03c8j are jointly constrained; see Section V for more details. We assume that all agents move with the unit forward speed. It is then straightforward to show that the angles \u03c6j satisfy the differential equations \u03c6\u0307j = \u2212 1 lj+1 (sin\u03b1j + sin(\u03c8j+1 + \u03b1j+1)) The angles \u03c8j of the polygon describing the formation satisfy \u03c8j = \u03c0 + \u03c6j \u2212 \u03c6j\u22121 and therefore \u03c8\u0307j = \u2212 1 lj+1 (sin\u03b1j + sin(\u03c8j+1 + \u03b1j+1)) + 1 lj (sin\u03b1j\u22121 + sin(\u03c8j + \u03b1j)) (1) For the distances lj one can similarly derive the equations l\u0307j = \u2212(cos\u03b1j\u22121 + cos(\u03c8j + \u03b1j)) (2) (cf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003418_j.aam.2020.102083-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003418_j.aam.2020.102083-Figure3-1.png", + "caption": "Fig. 3. The graph of H, H\u030c, H\u2217 and H\u030c\u2217 and the figures of four curved foldings {Pi(\u03b5)}i=1,...,4 for \u03b5 = 1/5.", + "texts": [ + " P1(\u03b5)), and also P3(\u03b4) (resp. P4(\u03b4)) cannot be a subset of P4(\u03b5) (resp. P3(\u03b5)). So we may assume that i = 1, 2 and j = 3, 4 without loss of generality. Since T (Pi(\u03b4)) \u2282 Pj(\u03b5), we have T (C0) = C0. Since C0 has no symmetry, T must be the identity map. So we have Pi(\u03b4) \u2282 Pj(\u03b5). To make it easy to see the subscriptions, we set (cf. (1.8)) H(t) := |Hf (t)|, H\u030c(t) := |Hf\u030c (t)|, H\u2217(t) := |Hf\u2217(1 \u2212 t)|, H\u030c\u2217(t) := |Hf\u030c\u2217 (1 \u2212 t)|. Using (1.8), one can compute that H, H\u030c, H\u2217 and H\u030c\u2217 and can plot the graphs as indicated in Fig. 3 (left). Then we can observe the relation H\u2217 < H\u030c < H < H\u030c\u2217 hold on the interval [1/10, 9/10]. Thus, the images of the four surfaces are mutually distinct. So we obtain the conclusion. In Fig. 3 (right), the images of the four surfaces f, f\u030c , f\u2217 and f\u030c\u2217 are indicated, which are all distinct. In Fig. 4, the crease patterns with ruling directions of the four curved foldings {Pi(\u03b5)}i=1,...,4 for \u03b5 = 1/5 are drawn. The difference of ruling angles in these four patterns also implies that the four curved foldings have distinct images (the noncongruence of them does not follow from this fact directly). \u2022 If C or \u0393 has a symmetry, the associated four curved foldings may not be noncongruent. For example, we set I := [\u2212\u03c0/6, \u03c0/6] and consider a part of the unit circle c1(t) = (cos t, sin t, 0) (t \u2208 I) and set C1 := c1(I)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001301_s11665-015-1511-4-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001301_s11665-015-1511-4-Figure9-1.png", + "caption": "Fig. 9 Temperature history ( C) of transverse cross section of weld calculated using cross section information given in Table 2 for solidification boundary, where Dt = Dl/V, Dl = (12.7/60), mm and V = 5.0 mm/s (Weld 2)", + "texts": [], + "surrounding_texts": [ + "Although the calculations presented here demonstrate certain aspects of the methodology, there still remain further investigations concerning its practical implementation, which would be in terms of algorithm development. It is therefore important to examine aspects of the inverse analysis methodology that are relevant to its further development and refinement. The general procedure for inverse thermal analysis of welds as described in this study includes interpolation between constrained isothermal boundaries, e.g., TTB and TM. A specific procedure for interpolation, however, has not been considered. Specific procedures for interpolation between constrained Journal of Materials Engineering and Performance surfaces within a temperature field are therefore an issue for algorithm development. Accordingly, further investigation is needed to determine a general and optimal procedure for interpolation between constrained isothermal surfaces in three dimensions. Toward this goal, the formal mathematical foundation of interpolation between constrained boundaries and various paths for linear interpolation is discussed. Formally, a solution to the steady-state advection-diffusion equation V @T @x \u00bc r \u00f0jrT\u00de; \u00f0Eq 6\u00de is to be determined within the region of the temperature field bounded by isothermal boundaries at TTB and TM, and surface and symmetry plane boundaries, which are nonconducting. This region of the temperature field is such that Eq 6 can be solved sequentially within one-, two-, and three-dimensional subdomains using direct model-based numerical procedures. These subdomains and the associated partitioning of Eq 6 are shown in Fig. 14 for the calculated temperature field of Weld 2. Following an inverse analysis approach, however, a solution to Eq 6 can be calculated Journal of Materials Engineering and Performance using interpolation procedures, which can be applied owing to the close proximity of boundary conditions, in contrast to more complex direct model-based procedures. The numerical solution of Eq 6, for specified boundary conditions, which is by application of various numerical algorithms (e.g., methods based on finite differences and volumes), is of the general form Journal of Materials Engineering and Performance Tn\u00fe1 p \u00bc P6 k\u00bc1 wkTn k P6 k\u00bc1 wk ; \u00f0Eq 7\u00de where the weight coefficients wk and iteration sequences, indexed by n, are specified according to a given solution algorithm. The grid indices k specify nearest neighbor grid locations relative to the grid location specified by index p. The general form of Eq 7, with constraints conditions specified according to Eq 3b, is that of the discrete interpolation operator. Accordingly, an important property following from this form is that for temperature field regions between boundaries, which are closed and in close proximity, regardless of the type of numerical procedure, the numerical solution of Eq 6 is equivalent to interpolation between boundary values. A heuristic discussion of this property, which is associated with the transition of numerical solvers for parabolic equations to those for elliptic equations, according to boundary conditions, is given in reference (Ref 24) with respect to finite volume discretization. Examination of the temperature field boundary conditions associated with both workpiece geometry and constrained isothermal surfaces TM and TTB establishes various paths for linear interpolation (see Fig. 14). Following these paths, interpolation between constrained surfaces can in principle be partitioned into a series of one-, two-, and three-dimensional interpolations. Referring to Fig. 14, one observes that boundary conditions on the calculated temperature field are such that interpolation can be applied in one dimension along the symmetry line at the top surface, in two dimensions over the top surface and symmetry planes, and in three dimensions over the closed domain bounded by isothermal surfaces TM and TTB, and the top surface and symmetry planes. As discussed previously (Ref 2), the embedding of constrained isothermal surfaces, which are constructed according to experimental measurements, within the calculated temperature field, compensates for errors due to estimated values of the diffusivity function, which are in fact a function of temperature. This follows in that embedded isothermal surfaces, which are sufficiently distributed volumetrically, represent implicitly the dependence of diffusivity on temperature, as well as advective influences within the melt pool. Multiple constrained isothermal surfaces can be constructed using thermocouple measurements following the same procedure applied in this study. Journal of Materials Engineering and Performance Parametric temperature histories obtained by inversion can be adopted for construction of databases, which in turn can provide for optimal inverse analysis of welding processes spanning a wide range of process parameters. The concept of constructing such databases is well posed in that, although there exist many different types of welding processes and their potential modifications, the physical characteristics of all types of welds are within a limited range of variation. That is to say, for all types of welding processes, the range of variations in solidification boundary shapes (and associated isothermal surfaces), and in weld thickness, is finite. In addition, for all metals and their alloys, the range in variation of material properties is also finite. It follows that a database of parametric temperature histories can be constructed that spans a wide range of welding processes, as well as their potential modifications. The availability of a parametric temperature history database can provide both parameter estimates for inverse analysis and correlation with documented weld analyses, both experimental and computational. The relationship between such a database and other components of the general framework for inverse thermal analysis using constraints is shown in Fig. 15. Referring to Fig. 15, it is seen that the results of both laboratory and computational experiments provide constraint conditions for inverse thermal analysis, while basic theory (heat transfer due to advection and diffusion) provides formulations of parametric models, i.e., different types of basis functions. In addition, the general framework shown in Fig. 15, which describes the use of constraint conditions obtained from computational experiments, establishes a well-defined relationship between direct and inverse modeling. An elucidation of this relationship is of significance in that many times direct and inverse models are compared unfairly. A discussion of such comparisons is beyond the scope of this study. It should be noted, however, that because of the interrelationship between direct and inverse models, it should be difficult to establish common criteria for their general comparison. In general, both direct and inverse models have aspects that contribute to computational cost, especially with respect to optimum use of cutting edge computational resources. As discussed above, the computational cost of direct models follows naturally from detailed mathematical representation of underlying physical processes. In contrast, the computational cost of inverse models follows from the ill-posedness of inverse problems, whose solutions are not unique. In particular, the inverse heat conduction problem (IHCP) is generally ill-posed, and thus methods of problem regularization are required, which tend to be iterative and computationally intensive (Ref 16). The \u2018\u2018inverse weld analysis problem\u2019\u2019 as specified in this study, although related to the IHCP, is well posed and not computationally intensive, as demonstrated by this and other studies (Ref 1, 2). References (Ref 16, 25) present discussion of illposedness of inverse problems associated with the IHCP and engineering applications, in general. The inclusion of a parametric temperature history database within the general inverse analysis framework (see Fig. 15) provides, in addition to parameter estimates, correlations with other type of weld analyses, such as microstructural analysis, as well as with weld processes and associated process parameters that are both feasible and optimal, which are documented in the literature. Finally, the present study applies a parametric model that is in terms of numerical-analytical basis functions, whose parameters are conveniently adjustable with respect to multiple constraint conditions. This parametric model is an inverse model formulation at its foundation. Direct model formulations, however, can also be applied parametrically for inverse analysis. For example, references (Ref 11, 26-28) present inverse analyses using direct models, and associated numerical simulations, for the determination of unknown variables. An issue to be addressed for inverse analysis using direct models is the feasibility of parameter adjustment with respect to multiple constraint conditions." + ] + }, + { + "image_filename": "designv11_34_0001793_taai.2015.7407100-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001793_taai.2015.7407100-Figure2-1.png", + "caption": "Fig. 2. Four screenshots from the simulator with the agent following different decision paths", + "texts": [ + " Further, compared to the architectures above, at the moment a purely case-based decision-making is implemented and not a heuristic search through rules. Finally, SiMA is based on psychoanalysis. There are doubts about whether it is a valid approach or not. An argument for a valid approach is that SiMA has defined most of the components of other architectures e.g. an attentional mechanism. These concepts were directly derived from the theory. IV. USE CASE DESCRIPTION The desciption of the decision-making is supported by a simple use case. It shall be used to clarity the concepts presented and at the same time, it is a part of the results. In Fig. 2, illustration 1, three objects are present: An agent at the bottom of the illustration controlled by SiMA decision-making, a food source for the top and a stone between the agent and the food source. The agent has a simulated body with homeostasis much like our own. At the start of the scenario, the main goal of the agent is to find a food source to eat from, i.e. to satisfy the \"hunger\"-drive. There are other drives as well as a \"relax\"-drive and \"deposit\" drive to excrement consumed food. The food source is visible to the agent", + " [f there were a hostile agent in the use case, there would be an episode that tells the agent about an experience with the hostile agent. The internal state of the agent would be stored as a memorized feeling. Options that originates from feelings tell the agent to react to external stimuli. This is a major difference to drives, which only fulfill bodily needs and cannot be used to avoid something that has been perceived. [n this use case, three options are extracted, and their applications are shown in Fig. 2. First, in \"2)\", an option is extracted from an episode about how to go around a stone. Second in \"3)\" an option is extracted from a drive, in order to make it possible to search for the desired percept. Third in \"4)\", an option is extracted from the percept \"Cake\", which would fulfIll the \"hunger\"-drive. [n 1 b) of Fig. 3, the current state of feelings is extracted from the emotions. Feelings are a description of the current mood and influence the selection of options. Each feeling stands for a certain preference behavior", + " It allows the system to consider new external influence in each cognitive cycle as new options are always evaluated together with the currently processed options. VI. RESULTS Decision-making of the SiMA agent was implemented in JA VA as an agent in the MASON framework [5]. The use case described is very simple, yet it demonstrates the basic functionality of decision-making in SiMA. The use case itself is a part of the results. Table 1 is a part of the simulator results, and it shows which internal actions are necessary to execute, in order to perform an external action. In Fig. 2, illustration 2), as the agent gets to the stone, \"cycle\" 13 is reached in Table I. Because the content of an episode perfectly matches the perceived state, the option to satisfy the \"hunger\"-drive is selected for this episode (\"ACTDRIVE\" in Table 1). The episode tells the agent how to get around the stone. In illustration 3) of Fig. 2, the agent was manually put somewhere else on the map. Now, no options were available from the perceived state or from any episode. Instead, the selected option originated from the \"hunger\"-drive. It triggered a search behavior. In illustration 4) of Fig. 2, after searching a while, the \"Cake\" is visible once again. From there, the option that originates from the perceived state (\"PERCEPTDRIVE\" in Table 1) is selected, because it is better to go to a \"Cake\" in the vision than to search for another one. The shown use case was a very simple one, but the purpose was to demonstrate the decision-making process. More impressing are the results of [ I], [5] and [16], which all rely on this decision-making functionality. VII. CONCLUSION The purpose of the project SiMA is to create a model of the human mental apparatus" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002971_j.matpr.2020.02.222-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002971_j.matpr.2020.02.222-Figure1-1.png", + "caption": "Fig. 1. Coaxial nozzle.", + "texts": [ + " Table 1 shows the chemical composition of Colmonoy 52SA powder used in this study. Stainless steel 316 was used as the substrate material for directed energy deposition of Colmonoy 52SA. The composition of stainless steel 316 is 12% (Ni), 17% (Cr), 2.5% (Mo), 1% (Si), 2% (Mn), 0.08% (C), 0.045% (P), 0.03% (S) and remaining is Fe. Colmonoy 52SA powder was deposited using DMD 105D machine with 1 kW diode laser. The laser has a wavelength of 940 nm and beam diameter of 0.3 mm. DMD 105D machine consists of coaxial nozzle where the powder is fed coaxially to the laser beam. Fig. 1 shows the cross-section of coaxial nozzle. It consists of three streams which impinge on the solid surface. These are nozzle gas, powder stream and shielding/shaping gas. Mixture of argon and helium was used as shield gas; argon was used as shaping gas which also creates an inert atmosphere to avoid possible oxidation. B C Cr Fe Si Ni 2.2% 0.55% 12.3% 3.8% 3.7% Balance Please cite this article as: B. N. Manjunath, A. R. Vinod, K. Abhinav et al., Optim energy deposition process, Materials Today: Proceedings, https://doi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure3-1.png", + "caption": "Fig. 3. Snakeboard motion.", + "texts": [ + " The Snakeboard shares the idea of forward propelling without kicking the ground. This system consists of a longitudinal bar that connects two sets of wheels, being allowed to rotate independently. In order to achieve a forward motion, the rider carries out alternate rotations with their ankles (exerting torques r and f on the rear and front wheelsets, respectively) and torso (torque \u03c8 ). These rotations are illustrated with the angles \u03c6r , \u03c6f and \u03c8 in Fig. 2 . In this way, the sequence of motion is shown in Fig. 3 . Firstly, both platforms are turned in, and the twist of the rider\u2019s torso through an angle \u03c8 allows the motion depicted by configuration (1) in Fig. 3 . Subsequently, the rider turns both feet out and moves the torso in the opposite direction, describing the motion denoted by (2). Repetition of this sequence over time leads to the forward propelling of the Snakeboard. In the case of the Waveboard, the forward propelling results from the oscillatory motion of both decks through lateral actions exerted by the rider with their feet. In this way, the lateral forces press the platforms down, making them roll about the axis of the Waveboard, and each pasive caster wheel is pushed out to the opposite direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001308_iemdc.2013.6556236-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001308_iemdc.2013.6556236-Figure3-1.png", + "caption": "Fig. 3. Mesh and results of stress analysis.", + "texts": [ + " The stress in the core is estimated by well-known 2-D plane stress analysis, which equations are as follows: 0 yx xyx (1) 0 yx yxy (2) where x and y are the x and y components of the stress, respectively; xy is the shearing stress, x, y, and xy can be expressed by the derivatives of the displacement at each node of the finite elements. Then, these equations can be solved by regarding the displacement as unknowns. Finally, the principal and Mises stresses at each finite element can be calculated, as follows: 2 2 1 42 1 xy yx yx (3) 2 2 2 42 1 xy yx yx (4) 2 21 2 2 2 12 1 VM (5) where 1 and 2 are the maximum and minimum principal stresses along the principal axes, respectively; VM is the Mises stress. Fig. 3 shows the finite element mesh used for the stress analysis and the calculated stress distributions. In this analysis, the nodes on the surface of the stator core are assumed to be displaced 25 m inside by the shrink fitting. The computational time of this analysis is 4 seconds by using Core i7 2.8 GHz PC. It is observed that the circumferential compressive stress is generated at the stator yoke by the shrink fitting. This stress causes the increase in the reluctivity and core loss of the stator core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000230_ijcaet.2017.086921-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000230_ijcaet.2017.086921-Figure5-1.png", + "caption": "Figure 5 Finite element model of inner and composite spring showing excitation during modal analysis (see online version for colours)", + "texts": [ + " Since each of the eigenvectors cannot be null vectors, the equation which must be solved is, ( )[K] \u03bb[M] {0}\u2212 = (6) For the modal analysis of inner and composite primary suspension spring in finite element analysis tool, ANSYS, the imported geometry is fixed at the bottom end in all degree of freedom and allow vertical deformation on the top face and solve for the possible ten modes gives natural frequencies shown in Figure 6. A three-dimensional model of inner and composite spring prepares in modelling software pro-engineer and it is imported in IGES file format in ANSYS for analysis. The imported geometry is mesh with solid 187 ten-node tetrahedral element with fine meshing. For analysis purpose structural steel plates are placed at bottom and top surface of spring of material chromium vanadium. The excited model of inner and composite spring using finite element analysis is shown in Figure 5 and ten modes of natural frequencies in Hz is shown in Figure 6. The theoretical natural frequency of inner spring is 38.06 Hz calculated by using following relation. n 2 d 6Ggf 2\u03c0D n \u03c1 = (7) In a structural system, any sustained cyclic load will produce a sustained cyclic or harmonic response. Harmonic analysis results are used to determine the steady-state response of a linear structure to loads that vary harmonically with time, thus enabling to verify whether or not the designs will successfully overcome resonance, fatigue, and other harmful effects of forced vibrations (Triana and Fajardo, 2013)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002344_s00170-016-9132-0-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002344_s00170-016-9132-0-Figure1-1.png", + "caption": "Fig. 1 Infrared staking concept [1]", + "texts": [ + " Until recently, traditional staking methods based upon the use of hot air, ultrasonication, or heated tools were widely used in a variety of industrial applications for joining plastic materials. Unfortunately, these methods have weaknesses, such as inducing stress in the formed stud, damage to components, plastic stringing or sticking to the punches, and inconsistent processing. Therefore, the development of new staking processes is desirable. The concept of infrared staking (IS) technology, which can be considered a new joining technique to assemble different plastic parts, is illustrated in Fig. 1 [1]. The IS process involves four basic phases: clamping, heating, forming, and cooling. During the IS process, the stud is heated evenly by infrared irradiation from a halogen lamp. At the end of the heating cycle, the semi-molten stud is then formed by a staking punch. Subsequently, ambient air is used to assist in regulating the punch and stud temperature. At the end of the hold time, the staking punch is retracted from the molded part, completing the cycle. This advanced process offers some advantages compared to traditional staking methods, such as higher efficiency and productivity and better mechanical properties", + " The geometric models of the workpiece and staking punch were generated using CATIA V5R20 and then transferred to DEFORM 3D by means of an STL-format file. The thermal-physical properties of the workpiece and staking punch were assumed to be constant and are presented in Table 2. Flow stress curves of the PP material changed as a function of the strain, strain rate, and temperature, as shown in Fig. 7 [14]. The following law was used owing to its ability to model the true behavior of a material: \u03c3 \u00bc \u03b5; \u03b5; T \u00f01\u00de where \u03c3, \u03b5, \u03b5, and T are the flow stress, effective plastic strain, effective strain rate, and temperature, respectively. As shown in Fig. 1, during the first stage (the heating stage), radiant energy generated by a halogen lamp was used to increase the plastic temperature. The lamp was composed of a coiled tungsten filament contained in a bulb enclosure filled with argon gas. Moreover, the lamp was coated with a reflector to increase the heat flux received by the product. A schematic representation of the halogen lamp, as shown in Fig. 8a, was used to calculate the filament and bulb temperatures based on the following nonlinear system of equations [16]: P\u2212\u03b5F T F\u00f0 \u00deS F\u03c3T 4 F\u22122\u03c0LBkArgon T* T F\u2212TB ln dB=d F\u00f0 \u00de \u00bc 0 \u00f02\u00de \u03b1B T F\u00f0 \u00de\u03b5F T F\u00f0 \u00deS F\u03c3T 4 F \u00fe 2\u03c0LBkArgon T* T F\u2212TB ln dB=dF\u00f0 \u00de \u2212\u03b5BSB\u03c3T4 B\u2212hSB TB\u2212T\u00f0 \u00de \u00bc 0 \u00f03\u00de where P, \u03b5, S, L, and d represent the lamp power, emissivity, area, length, and diameter, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000408_ccece.2015.7129153-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000408_ccece.2015.7129153-Figure2-1.png", + "caption": "Fig. 2. Maxwell Coils. A Maxwell coil is a device for producing a large volume of almost constant (or constant gradient) magnetic field. A Maxwell coil consists of a pair of coils separated by \u221a 3 times their radius. Each coil carries an equal electrical current flowing in the different direction.", + "texts": [ + " Compared with the permanent magnet, the electromagnetic coils are more flexible to control and alter the magnetic field. Therefore electromagnetic coils are widely used for actuation. There are two types actuation force that can be used for remote manipulation, the magnetic torque and the magnetic force. Magnetic torque makes the magnet rotating and the magnetic force pull or push the magnet. There are two kinds of coils based actuation methods, 1) the uniform magnetic field and constant gradient magnetic field, which are generated by Helmholtz coils (as shown in Fig.1) and Maxwell coils (as shown in Fig.2), respectively [6]\u2013[9]. The Helmholtz and Maxwell coils usually work together to control a micro magnetic particle. The model of these two coils are very simple and therefore the system is easy to be realized. The uniform field from Helmholtz coils provide the torque and the constant gradient from Maxwell \u2217Corresponding authors are Ren H. and Yu H., mail:{ren, bieyhy}@nus.edu.sg This work is supported by Singapore Academic Research Fund, under Grant R397000139133, R397000157112, R397000156112 and NUS Grant C397000039001" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002999_1464419320910864-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002999_1464419320910864-Figure1-1.png", + "caption": "Figure 1. (a) Geometric representation of roller bearing (b) geometrical deformation between the roller and races.", + "texts": [ + " A perfect rolling motion is considered between rollers and raceways, (both inner race and outer race) and there are no slipping, skidding or skewing motion involved in the system. . The elastic contact deformation between the roller and raceways satisfies the Hertzian contact theory. . The bearing operates under the isothermal conditions. Only the hysteretic or material damping is considered, neglecting the effect of lubrication in the contact areas. A set of nonlinear equations, using the NR\u00fe 2 independent coordinates are formed from the Lagrange\u2019s equation. From Figure 1, the geometrical deformations between rollers and races can be expressed as ir \u00bc r\u00fe Rr \u00fe cl\u00f0 \u00de j or \u00bc R j \u00fe Rr \u00fe cl \u00f01\u00de As the deformations at the contact points are defined using the Hertzian contact theory, both ir and or are effective when the springs are compressed due to contact as shown in Figure 1(b). Hence, the contact force between roller and raceways using the contact deformation theory can be written as Fj \u00bc k 10=9j U\u00f0t\u00de \u00bc Kj j U\u00f0t\u00de \u00f02\u00de where U\u00f0t\u00de; 1, If j 4 0 0, If j 5 0 In the case of a roller bearing, an elliptical patch is formed at the contact points of the rollers and raceways, which is calculated by the Hertzian contact theory. The profile of the roller has significance while considering the deformation. In the present study, the profile of the roller has been considered to be cylindrical", + " Using Lagrange\u2019s theory, the contribution of each element of roller bearing system in the form of kinetic and potential energy is formulated in the form of equations of motion with respect to the generalized coordinates j, xir and yir mr \u20ac j\u00femr j _ 2 j \u00femrgsin j\u00fe 1 2 @ Kir 10=9 1 ir h i \u00fe @ j ir\u00bd 2 \u00fe Kir 10=9 ir h i \u00fe @ j @ j \u00fe 1 2 @ Kor 10=9 1 or \u00fe h i @ j or\u00bd 2 \u00fe \u00fe Kor 10=9 or \u00fe h i \u00fe 10 9 XNR: j\u00bc1 Cir Kir\u00f0 \u00de 10=9 ir\u00fe j q @ _x @ _ j \u00fe 10 9 XNR: j\u00bc1 Cor Kor\u00f0 \u00de 10=9or\u00fe _ j q \u00bc 0 \u00f06\u00de mir \u20acxir XNR j\u00bc1 Kir 10=9 ir h i \u00fe @ j @xir \u00fe 10 9 XNR j\u00bc1 CirKir 10=9 ir h i \u00fe _xj q @ _xj @ _xir \u00bc Fu sin! t \u00f07\u00de mir \u20acyir \u00femirg XNR j\u00bc1 Kor 10=9 or \u00fe h i @ j @yir \u00fe 10 9 XNR j\u00bc1 CirKir 10=9 ir h i \u00fe _xj q @ _xj @ _yir \u00bcW\u00fe Fu cos! t \u00f08\u00de From Figure 1, Xj the distance from the inner race center to the roller center can be calculated as j \u00bc xor xir\u00f0 \u00de 2 \u00feR2 r \u00fe 2Rr xor xir\u00f0 \u00de cos j \u00fe 2Rr yor yir\u00f0 \u00desin j \u00fe yor yir\u00f0 \u00de 2 \" # \u00f09\u00de @ j @ j \u00bc j \u00fe xor xir\u00f0 \u00de cos j \u00fe yor yir\u00f0 \u00desin j j @ j @xir \u00bc xor xir\u00f0 \u00de j cos j j @ j @yir \u00bc yor yir\u00f0 \u00de j sin j j @ _xj @ _xir \u00bc xir xor\u00f0 \u00de j cos j j @ _xj @ _yir \u00bc yir yor\u00f0 \u00de j sin j j The above set of equations (6) to (8) are nonlinear (NR \u00fe 2), second-order differential equations. Rotor mass and its unbalanced effect have been added and no other external load interaction with the system has been considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001232_isic.2014.6967614-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001232_isic.2014.6967614-Figure4-1.png", + "caption": "Fig. 4: Graph representation of Configuration 1 and 2.", + "texts": [ + " PyZMQ is a Python binding for the networking library \u00d8MQ, an emerging solution for distributed computing frameworks [19]. The serialization and compression of the NumPy arrays for the transmission was achieved by converting them into the MessagePack format. During the remainder of the paper, two configurations of the experimental setup with different coupling topologies will be presented to be controlled by the modular DMPC scheme developed so far. The coupling graph of Configuration 1 is depicted in Figure 4(a) where the control goal is to stabilize the water level in Tank 4 at the right hand side, subject to a constant water withdrawal. The corresponding agent is of the type sink. The control input of the system is the water injection of Tank 1 on the left hand side. Thus, the corresponding agent is of type source. The Tanks 2 and 3 are of the agent type link. For Configuration 2, an additional coupling between Tank 1 and 4 is introduced and Tank 4 is now equipped with a pump to inject water into the system while Tank 3 is subject to a constant water withdrawal by an outlet valve. Thus, Tank 1 and 4 are now source-agents, Tank 2 is of type link, and Tank 3 is a sink-agent. The coupling graph of Scenario 2 is shown in Figure 4(b). Both configurations exhibit the property that the graphs describing the agent couplings are not fully connected. Since the considered algorithm from Section II-B only utilizes neighbor to neighbor communication, this implies that not all agents communicate with each other and none of the agent possesses complete knowledge of the system. These are key characteristics of distributed systems, making the scenarios interesting test cases for DMPC. Furthermore the mutual state coupling of the tank dynamics leads to a more involved optimization problem compared to the input coupling encountered in the comparison of DMPC implementations running Johansson\u2019s process [9]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000493_asjc.933-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000493_asjc.933-Figure1-1.png", + "caption": "Fig. 1. The internal structure schematic of the morphing wing.", + "texts": [ + " This morphing structure utilized two degrees of freedom (2-DOF), the sweepback angle and the internal angle of the four-bar linkage Manuscript received July 27, 2012; revised October 13, 2012; accepted November 20, 2013. Rongqi Shi is with the Department of Engineering Mechanics, Tsinghua University, Beijing, China (corresponding author, e-mail: shirongqi@tsinghua. edu.cn). Jianmei Song is with the School of Aerospace, Beijing Institute of Technology, Beijing, China. Weiyu Wan is with the Beijing Electro-mechanical Engineering Institute, Beijing, China. assembly, to accommodate wing geometry changes, as shown in Fig.1. [3]. At present, research on the wing structure is mainly focused on the optimization of the actuator location. Sequence quadratic programming (SQP) is used to solve the optimization problem [4], and a two-stage optimization process based on the genetic algorithm and gradient-based optimization is developed for maximum energy efficiency [5]. NASTRAN and SQP are used to optimize the actuator location with the structure flexibility taken into account [6,7]. The research of this paper is mainly focused on the control system design of the morphing wing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003194_s12206-020-0503-y-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003194_s12206-020-0503-y-Figure4-1.png", + "caption": "Fig. 4. Collision model of the new-type TDB.", + "texts": [ + " (3) In the AMB system, the rotor is normally driven by an inverter to realize rotation. In industrial applications, the output of the inverter will be immediately cut off once AMB fails. Therefore, the rotor rotational motion equation around z axis after rotor drop for only rotor without load end and ignoring the wind friction can be written as r b1 b2\u03b8 = \u2212 \u2212zJ M M (4) where Mb1 and Mb2 are the frictional moments from the left and right TDB, respectively. Two coordinate systems (xoy and \u2032 \u2032x oy ) are established at the center of active TDB, as shown in Fig. 4, for ease of calculation. The collision force between the outer race of the ball bearing and the MSB 1 is taken as example. In the coordinate system xoy and \u2032 \u2032x oy of this example, the coordinates of the inner arc center of the MSB 1, respectively, refer to ( )h1 h1,x y and ( )h1 h1,\u2032 \u2032x y . The coordinates of the top edge of the MSB 1, respectively, refer to ( )h11 h11,x y and ( )h11 h11,\u2032 \u2032x y . The coordinates of the outer arc center of the outer race, respectively, refer to ( )b1 b1,x y and ( )b1 b1,\u2032 \u2032x y ", + " On the basis of Hunt\u2013Crossley collision theory, the collision force between the outer race and the MSB can be expressed as [20] 10/9 n h1 h1 h1 nh1 h1 3(1 ) >0 2 0 0 \u03b4 \u03b6\u03b4 \u03b4 \u03b4 \u23a7 +\u23aa= \u23a8 \u23aa \u2264\u23a9 K F (5) where nK refers to the contact stiffness. h1\u03b4 refers to the contact depth between the outer race and the MSB. The coefficient \u03b6 concerning the contact between steels is valued in the scope of 0.08\u20130.32 s/m; 0.2 s/m is chosen here. Palmgren\u2019s empirical formula is used to calculate the contact stiffness [21]: 10 8 9 n n8.06 10= \u00d7K L (6) where Ln refers to the contact width between the outer race and the MSB. As shown in Fig. 4, the coordinate ( )b1 b1,\u2032 \u2032x y of the outer race arc center can be used to judge that the outer race can only collide with the upper part of MSB 1. When calculating the contact depth, the collision direction should be considered, and the contact depth can be expressed as ( ) ( ) ( ) ( ) 2 2 b1 h1 b1 h1 h1 bo h1 2 2 b1 h11 b1 h11 bo \u03b3 \u03b1 \u03b4 \u03b3 \u03b1 \u23a7 \u2212 + \u2212 \u2212 + \u2264\u23aa= \u23a8 \u23aa \u2212 + \u2212 \u2212 >\u23a9 x x y y R R x x y y R (7) where \u03b1 refers to half of the arc angle of the MSB and \u03b3 is the angle between the collision direction and x\u2032 axis direction", + " (9) The coordinate value of the top edge of the MSB 1 in the co- ordinate system xoy can be expressed as ( ) ( ) h11 h1 h1 bo b0 1 h1 h11 h1 h1 h1 bo b0 1 2 cos sin 2 2 sin cos 2 \u03b1 \u03c2 \u03b1 \u03b1 \u03b1 \u03c2 \u23a7 = \u2212 \u2212 \u2212 + +\u23aa\u23aa \u23a8 \u23aa = \u2212 + + + \u2212\u23aa\u23a9 x R R R s R y R R R R s . (10) The calculation method of collision force between the outer race and MSB 2 is similar to the above-mentioned method. The position of the FSB remains unchanged. The collision force between the outer race of the ball bearing and the FSB can be expressed as [21] 10/9 n g1 g1 g1 ng1 g1 3 ) 0 2 0 0 \u03b4 \u03b6\u03b4 \u03b4 \u03b4 \u23a7 >\u23aa= \u23a8 \u23aa \u2264\u23a9 \uff081+K F . (11) As shown in Fig. 4, the coordinate ( )b1 b1,\u2032 \u2032x y of the outer race arc center can be used to judge that the outer ring can only collide with the left part of FSB 1. The contact depth can be expressed as ( ) ( ) 2 2 b1 b1 b0 g1 2 2 b1 g11 b1 g11 bo \u03b4 \u03d1 \u03b2 \u03b4 \u03d1 \u03b2 \u23a7 + \u2212 \u2264\u23aa= \u23a8 \u2032 \u2032 \u2032 \u2032\u23aa \u2212 + \u2212 \u2212 >\u23a9 x y x x y y R (12) where \u03b2 refers to half of the arc angle of the FSB and \u03d1 is the angle between the collision direction and y\u2032 axis direction. b1 b1 tan\u03d1 \u2032 = \u2032 x y (13) The coordinate value of the center of the outer race in the coordinate system \u2032 \u2032x oy can be expressed as ( ) ( ) b1 b1 b1 b1 b1 b1 2 2 2 2 \u23a7 \u2032 = \u2212\u23aa\u23aa \u23a8 \u23aa \u2032 = +\u23aa\u23a9 x x y y x y " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002755_1.5063139-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002755_1.5063139-Figure3-1.png", + "caption": "Figure 3: Failed Body centred cubic due to sudden increase in layer area", + "texts": [ + " Previous studies show that the SLM processing of aluminium and its alloys is more difficult than titanium or steel [12, 13]. On inclined surfaces there is a staircase effect which causes thermal heat concentration. These concentrations of thermal energy cannot easily conduct through the metal support powder and can result in warping and degraded surface quality (Figure 1 & Figure 2). Such overheating degrades the aesthetic and mechanical lattice properties, and in extreme cases can result in gross overheating (Figure 3). SLM research typically focuses on titanium alloys due to their desirable mechanical properties and biocompatibility, for example [14-18]. Aluminium has attracted less attention; however it is an important material for AM due to favourable mechanical properties, low cost, thermal conductivity and corrosion resistance [9, 19, 20]. There exists processing difficulties associated with aluminium powder, including poor flow-ability, high laser reflectivity and oxidation during melting [13]. Furthermore, aluminium has very high thermal diffusivity and melt-pool solidification can be a challenge for geometries with high heat input, and high resistance to thermal conductivity (Table 1)", + " The laser scans the metal powders at speed of 200-100mm/s causing an extremely high thermal gradient (typically in the range of 102 K/s) [6]. Components with sufficient material below the current layer can remove excess thermal energy by conduction. However, if a component\u2019s cross sectional area rapidly changes, high thermal concentrations can occur resulting in thermal overload. For example the increase in section of the Body Centred Cubic Structure of Figure 4 is hypothesised to have caused gross overheating due to the increase in heat input above a region of high thermal resistance (Figure 3). There are in excess of 130 parameters of influence to SLM manufacturability [6]. Furthermore, the SLM process is difficult to simulate due to: complex dynamic boundary conditions, including gas flow [21] and build plate heating [22]; multi-scaled features [23, 24]; high thermal gradients [25, 26]; uncertainty of powder material behaviour [27, 28]; and, complex meltpool dynamics[29], including three-dimensional thermo-capillary convective currents [30, 31]. Consequently, robust modelling of SLM manufacturability is extremely challenging" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003748_iecon43393.2020.9254887-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003748_iecon43393.2020.9254887-Figure2-1.png", + "caption": "Fig. 2. Vector diagram of IPMSM", + "texts": [], + "surrounding_texts": [ + "Cogging torque is an inherent property of permanent magnet motor, which is generated by the interaction between the magnetic poles of the rotor and the stator core. According to the definition of cogging torque, the following can be obtained: = \u2212 Where, is the cogging torque, is magnetic field energy, and is the relative position angle between stator and rotor. The magnetic field energy is given by: = \u222b From (5) , we can find that the magnetic field energy is determined by the structure and size of the motor, the relative position of the stator, and permanent magnet properties. The distribution of air gap magnetic density along the armature surface can be approximately expressed as: ( , ) = ( ) ( ) ( ) ( , ) where, ( ), \u03b4(\u03b8) and \u210e ( ) are remanence of permanent magnet, effective air gap length and the distribution of permanent magnet width along the circular direction respectively. Substitute (7) into (6), it becomes: = \u222b ( ) ( ) ( ) ( , ) Fourier decomposition is performed for ( ) and ( ) ( ) ( , ) respectively, and the result is expressed as: ( ) = + \u2211 2 ( ) ( ) ( , ) = + \u2211 where, is the number of stator slot. Substitute (9) and (10) into (8), and combine (5) to get the cogging torque expressed as (11), which is at the bottom of the page. where, is permanent magnet radius, is stator radius, is Fourier coefficient of flux density function, is Fourier coefficient of air gap permeance function, is the least common multiplier of stator slots and rotor poles. From (11), we can clearly find that, cogging torque can be reduced by modifying the Fourier series coefficient of air gap permeance function." + ] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure13-1.png", + "caption": "FIG. 13. Crease-bending model facet layout. A circumferential hinge is located at a distance from the central vertex. This layout deforms by rotation of the hinge lines only.", + "texts": [ + " We now modify this rigid-faceted approach to incorporate this observed behavior. B. Bending crease If bending is allowed along the crease line, then an alternative deformation mode characterized by a localized dimple around the indentation point occurs, as described in Sec. II. Following the rigid-faceted approach, we discretize each face 043001-7 between creases into a series of rigid facets and postulate a kinematic mechanism which captures the dimple deformation. The postulated facet layout is shown in Fig. 13. There are two key differences in comparison to the straight-crease model (Fig. 6). First, there is a circumferential hinge located at a radius of , measured along the crease from the center of 043001-8 the disk. Second, there are additional facets which emanate from the intersection of this hinge with the crease line (facets C and D). We assume that the facets directly adjacent to the crease (A\u2013F) have subtend angles of \u03b10, while all other facets have subtended angles of \u03b1 = (\u03c0 \u2212 2\u03b10)/(2n \u2212 1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002174_j.engfailanal.2016.06.014-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002174_j.engfailanal.2016.06.014-Figure14-1.png", + "caption": "Figure 14: Wheel gauge change as a function of train running time for two different braking histories.", + "texts": [ + " malfunctioning of distributor valves, clogging of brake pipe, etc.). Thus, non-uniformity may be higher for some braking cycles and lower later on or vice-versa. To investigate, the effect of varying non-uniformity in braking we consider two scenarios. First, we subject a wheel set to a braking cycle with maximum heat input factor 4 and then to a braking cycle with maximum heat input factor 2. Second, we subject a wheel set to a braking cycle with maximum heat input factor 2 and then to a braking cycle with maximum heat input factor 4 (see Fig. 14). It can be seen from the figure that final wheel gauge change at end of second braking cycle and maximum gauge reduction, are essentially identical for both braking scenarios suggesting that these are primarily determined by maximum heat input factor. It can be seen that gauge reduction is lower in subsequent braking cycles when heat input factor for subsequent braking cycle is lower than the earlier one. Fig. 15 shows residual hoop stresses for S-type coach wheels at the end of a braking cycle for maximum heat input factors one (synchronized braking) and four" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001297_j.jmatprotec.2013.01.009-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001297_j.jmatprotec.2013.01.009-Figure1-1.png", + "caption": "Fig. 1. Examples of analytical heat sources. (a) Goldak double ellip", + "texts": [ + " Critical review of heat sources Faced with the difficulty of developing an efficient global weldng simulation model, in terms of calculation time and simplicity f use, many authors have attempted to define the arc heat input n a simplified way. The method consists in applying a weld heat ource, the expression of which is defined analytically to represent he heat power delivered by the arc. Akbari Mousavi and Miresmaeili (2008) used a Gaussian surface eat flux, representative of the distribution of heat in the plasma sed in their study. For GTAW, Goldak et al. (1984) suggested using volumetric double ellipsoid shaped heat source, the parameters f which were calculated from experimental fusion zone measure- ents (Fig. 1). This weld heat source has also been used by several uthors to develop a numerical model for evaluating the effect of elding parameters on plate butt joint welding as studied by Gery t al. (2005) or to study the weldability of IN738LC superalloy like (Goldak et al., 1984). (b) Gaussian heat source (Depradeux, 2004). Danis et al. (2010). These methods imply an adjustment of the heat input parameters to make them fit with the experimental thermal field monitored during the tests. A trial was required prior the initiation of a welding simulation, it was, therefore, not predictive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003065_s40684-020-00217-3-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003065_s40684-020-00217-3-Figure1-1.png", + "caption": "Fig. 1 The five-axis hybrid machine tool with the laser module", + "texts": [ + " In addition, a laser-assisted fillet milling method was developed to machine the edges of a square member by combining the fillet milling method and laser preheating [23]. However, research on 3D shape machining using LAMill has been performed only for simple shape machining despite the necessity to process products with complicated shapes for commercialization. The purpose of this study is to fabricate a turbine blade using a five-axis hybrid machine tool with a laser module ultimately for the commercialization of LAMill. Figure\u00a01 shows a schematic diagram of the five-axis hybrid machine tool with the laser module. The laser module has two axes that can move in the x and y directions. The manufacturing of a final product using LAMill is attempted for the first time in this study. The shape of the turbine blade was modified for machinability in the currently constructed the five-axis hybrid machine tool with the laser module by reviewing several models actually used in industry. A thermal analysis using the modified 1 3 blade model was conducted to determine the cutting depth for LAMill" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001795_robio.2015.7418892-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001795_robio.2015.7418892-Figure9-1.png", + "caption": "Figure 9. (a) The building of arc coordinate system", + "texts": [ + " By reading the joint angles and solving the forward kinematic in real time, we can get the motion curve of terminal gripper in Cartesian space as shown in Fig 8. The purple curve is the ideal trajectory and the red one is the tested trajectory. We project the tested trajectory to X O Y plane and find that the maximum error is less than 0.6cm. Therefore, the manipulator shows well linear trajectory tracking and the trajectory planning method is effective. Attached the makers onto the door surface and their center position in base coordinate system can be detected. Then the arc coordinate system is built as shown in Fig 9. The processes of opening door are shown in Fig 10. The whole process can be divided into six steps. Firstly, the terminal gripper moves to the position which is away from the door handle about 15cm. Secondly, the terminal gripper moves close to the door handle. Thirdly, the gripper grasps the door handle. Fourthly, the gripper opens the door with arc motion. Then, the gripper releases the door handle. Finally, the gripper leaves the door handle. (b) the sketch of opening door. The door handle is fixed on a constant height, and the arc trajectory planning is just in X O Y plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002519_s40997-016-0052-2-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002519_s40997-016-0052-2-Figure2-1.png", + "caption": "Fig. 2 Schematic diagram of the linear two-mass system with backlash", + "texts": [ + " A two-mass system without backlash can be described with the following linear model (Nordin and Gutman 2002): Jm _xm \u00bc cmxm Ts \u00fe Tm Jl _xl \u00bc clxl \u00fe Ts Td hd \u00bc hm hl; xd \u00bc xm xl _hm \u00bc xm; _hl \u00bc xl; _hd \u00bc xd Ts \u00bc kshd \u00fe csxd \u00f01\u00de where hm and hl are motor and load angles, xm and xl are motor and load angular velocity, hd and xd are difference angle and difference angular velocity, Jm (kg m2) and cm (Nm/(rad/s)) are motor moment of inertia and viscous motor friction, and Tm (Nm) and Ts (Nm) are motor torque and the transmitted shaft torque, respectively. Jl (kg m2) and cl (Nm/(rad/s)) are load moment of inertia and viscous load friction, ks (Nm/rad) and cs (Nm/(rad/s)) are shaft elasticity and inner damping coefficient of the shaft, and Td (Nm) represents the load torque disturbance. In two-mass systems with backlash, the shaft is replaced by a shaft including backlash. A schematic diagram of a two-mass system with backlash is depicted in Fig. 2, where a is the maximum value of the backlash angle. Dynamical equations for a two-mass system with backlash are described as the following (Nordin and Gutman 2002): Jm _xm \u00bc cmxm Ts \u00fe Tm \u00f02\u00de Jl _xl \u00bc clxl \u00fe Ts Td \u00f03\u00de hd \u00bc hm hl; xd \u00bc xm xl \u00f04\u00de _hm \u00bc xm; _hl \u00bc xl; _hd \u00bc xd \u00f05\u00de Ts \u00bc kshs \u00fe csxs \u00f06\u00de _hs \u00bc xs \u00f07\u00de where hs denotes shaft twist angle andxs denotes shaft twist angular velocity. In two-mass systems without backlash, the shaft torque is related to the difference angle and its velocity, but in systems with backlash the shaft torque depends on the shaft twist angle and its velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001814_2016-01-1132-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001814_2016-01-1132-Figure1-1.png", + "caption": "Figure 1. Conventional Ball-type joint", + "texts": [], + "surrounding_texts": [ + "The powertrain of cars and light trucks has been evolving to reduce fuel consumption. Recent fuel efficiency improvements have been accomplished through various techniques including engine downsizing, turbocharging, stop-start systems and transmissions with more gears. Mechanical and friction losses in the engine still account for most of the mechanical energy losses in a vehicle, followed by losses in the transmission, aerodynamic drag and tire rolling resistance. Mechanical energy losses due to driveline and chassis components (i.e. differential, driveshafts, halfshafts, wheel bearings), in a typical compact car are in the order of 2% of the total mechanical energy losses [1]. Therefore, the contribution of the Halfshafts (HS) to fuel economy, with respect to other systems and sub-systems in the vehicle, is small. It is estimated that a 0.1 percentage point mechanical efficiency improvement in a set of Halfshafts would result in fuel economy improvements of approximately 0.1% or less, under typical highway conditions. Constant Velocity Joints (CVJ\u2019s) can be classified according to its architecture in two large groups: Ball-type joints (i.e. Rzeppa, Cross-Groove, Double-Offset) and Tripot-type joints (i.e. Standard and Premium Tripot). Relative motion, among the parts of a CVJ, has a rolling and a sliding component. Most of the frictional losses in CVJ\u2019s are caused by the sliding component. Therefore, if an improvement in mechanical efficiency is desired, two strategies (or the combination of both) can be followed which are: \u2022 Reduce the amount of sliding through kinematics, i.e. cage balancing through track-geometry [2]; \u2022 Reduce the frictional forces through a tribological approach (grease and texture). In this work, differences in mechanical efficiency among different types of CVJ\u2019s are quantified, as well as the improvement in mechanical efficiency that can be achieved through the right selection of grease for the application." + ] + }, + { + "image_filename": "designv11_34_0002010_lascas.2016.7451026-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002010_lascas.2016.7451026-Figure6-1.png", + "caption": "Fig. 6. Symmetrical structures with four variable capacitors. a) With synchronous movement. b) With complementary movement.", + "texts": [ + " The connections for energy harvesting are the same, with the negative output used just as bias. IV. GENERATORS WITH FOUR VARIABLE CAPACITORS The structures in Fig. 5 can also be operated with C1 and C2 varying, forming complementary pairs with Ca and Cb, what can be easily achieved with the experimental structure by taking the outputs at the moveable plates and grounding the two unused lower combs. These structures require less minimum capacitance variation, of 1:1.414 for Fig, 6a and 1:1.618 for Fig. 6b. They recover faster than the others do, especially Fig. 6b, what allows a significant increase in the harvested energy. They do not appear to correspond to previously proposed structures. Other connections are also possible, as connecting C1 and C2 in parallel between nodes 1 and 4 in Fig. 6a, what results in some improvement. It is possible to generate two outputs with the same polarity by inverting D1-D4 or D2-D3 in Fig. 6b, and in this case Ca and Cb can be fixed, in a symmetrical version of Fig. 3b. ISBN 978-1-4673-7835-2/16/$31.00\u00a92016 IEEE 128 IEEE Catalog Number CFP16LAS-ART V. SIMULATIONS AND EXPERIMENTAL EVALUATION The structures were evaluated connected to the same DC/DC converter described in [3], represented in a simplified way in Fig. 7, with a generator as the one in Fig. 3a. The converter senses the output voltage at node 1 through a capacitive divider C11-C12, and when it reaches a certain level briefly switches on the transistors M1 and M2, operating the buck converter and resetting the capacitive divider" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001225_dscc2014-6114-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001225_dscc2014-6114-Figure4-1.png", + "caption": "Figure 4. Terminal equality constraint", + "texts": [ + " x\u0304s(k+N) \u2208 Reach(x\u0304s(k+N\u22121)) (11) where x\u0304s is the reference state for the coarse stage and Reach is defined as Reach(X ) = {x+ \u2208 Rn : \u2203x \u2208 X , \u2203u \u2208U s.t. x+ = Ax+Bu} (12) To avoid numerical instabilities and enlarge the domain of feasibility of the problem, the terminal constraint is commonly expressed as an inequality constraint. Therefore, the reachability condition (11) can be formulated as Pre ( X\u0304s(k+N) ) \u2229 X\u0304s(k+N\u22121) = X\u0304s(k+N\u22121) (13) which can be more efficiently computed, and where the terminal set is defined as X\u0304s = {x \u2208 Rn :\u2212[\u03b4pos \u03b4vel] \u2032 \u2264 x\u2212 x\u0304s(k)\u2264+[\u03b4pos \u03b4vel] \u2032} (14) and is shown in Fig. 4 and Pre is defined as Pre(X ) = {x \u2208 Rn : \u2203u \u2208U s.t. Ax+Bu \u2208 X } (15) Further details on Pre and Reach sets can be found in [12]. Computation of Pre-set for a controlled system involves set projection. Here, we propose a different approach for evaluating condition (13). Proposition 1. Given x(k+1) = Ax(k)+Bu(k), a polytope of admissible input U, a polytope X with vertices V = {v1, \u00b7 \u00b7 \u00b7 , ve}, and a polytope X +, if for every vi, i = 1, \u00b7 \u00b7 \u00b7 , e, there exists ui \u2208 U such that Avi +Bui \u2208 X +, then for every x \u2208 X , there exists u \u2208U such that Ax+Bu \u2208 X + The proof of Proposition 1 is given in Appendix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002042_tmag.2016.2527059-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002042_tmag.2016.2527059-Figure1-1.png", + "caption": "Fig. 1: Picture of a bearingless permanent magnet machine prototype (a) and schematic representation (b).", + "texts": [ + " This eliminates the need for mechanical bearings and allows the rotor to be operated in a sealed compartment. Bearingless machines are well suited for applications with demands on high speeds, low wear and low particle generation and contamination [1]\u2013[3]. This paper focuses on a bearingless permanent magnet disc drive [4]. The rotor consists of a ring-shaped permanent magnet with a diameter that exceeds its axial length. This topology is advantageous because it features passive stability in three degrees of freedom. Therefore, only three degrees of freedom have to be controlled actively [5], [6]. Figure 1 shows a bearingless machine that utilizes such a topology and is used to verify the angle estimation method. The control system utilizes information from displacement and angular sensors to achieve stable levitation and a high drive performance. The cost and complexity of the system can be reduced if an angle-sensorless control is used. This allows for new areas of operation for bearingless machines. Different ways of sensorless estimation of the rotor angle for permanent magnet machines are reported in the literature [7]\u2013[10]", + " Figure 4(b) shows the same example with an angle estimation error \u2206\u03b8 6= 0. The net force acting upon the rotor is still zero, therefore ~FB is opposing ~Fr,e. However, F \u2217 B and subsequently ~IB are rotated by \u2206\u03b8 according to (8). The angle observer detects that the angle between ~Fr,e and ~IB is no longer 180 \u25e6 and corrects the angle estimate \u03b8\u0302 to drive \u2206\u03b8 to zero. Figure 5 shows a simulation of the operation of a bearingless machine operated with the angle estimator. The simulation models the machine shown in Fig. 1 and captures its dynamics by including all flux density harmonics listed in Tab. II. Futher non idealities such as measurement noise are also modelled. The simulation assumes that there is no radial gravitaional pull and therefore, the radial reference position is set to r\u2217 = \u221a (x\u2217) 2 + (y\u2217) 2 > 0. (12) This results in a radial displacement force ~F\u2206 according to (1) 0018-9464 (c) 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee", + " The integrator in (11) drives the estimation error \u2206\u03b8 to zero once the rotor speed reaches the reference speed, c.f. Fig. 5(b). The estimation error is shown in Fig. 5(b). The rotor speed estimate is calculated by derivation of the angle estimate and is used for the drive current controller. Figure 5(a) shows the machine speed n, the reference speed n\u2217 and the machine speed estimate n\u0302. The machine speed is constant after approximately 11 s. V. IMPLEMENTATION The angle estimation method has been implemented on a commercialy available machine, cf. Fig 1(a) to verify the functionality of the model based angle estimation. The control utilizes only the estimated angle for operation, angle sensors are used to calculate the estimation error. The machine speed is obtained by taking the time derivative of the estimated angle. The angle-sensorless initial angle detection and start-up procedure are taken from [11]. The novel angle estimation method is started once the rotor reached stable levitation. The rotor refrence position was displaced to generate a sufficient external radial force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001027_detc2013-13712-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001027_detc2013-13712-Figure2-1.png", + "caption": "Figure 2. Cross section of tiling-pad bearing", + "texts": [ + " NOMENCLATURE B1, B2 Weight Factors of Axial Pressure Profile Cs Linear Rotor Damping Matrix Cb Bearing Radial Clearance (m) Cp Pad Radial Clearance (m) D Bearing Journal Diameter (m) Fbrg Bearing Forces (N) Fg Gravitational Force Vector Fn Nonlinear Force Vector Fp Pad Force (N) Fu Unbalance Force Vector Fx, Fy Bearing Force in x, y direction (N) G Gyroscopic Matrix h Fluid Film Thickness (m) h Dimensionless Film Thickness ( pCh / ) Ipad Matrix of Pad Moments of Inertia (kg-m2) Kb Stiffness Matrix of Bearing Kpiv Pivot Stiffness (N/m) Ks Stiffness Matrix of Shaft L Bearing/Damper Length (m) m Preload of Tilting Pad Bearing ( pb CC /1 ) Mb Mass Matrix of Bearing Ms Linear Rotor Mass Matrix Mp Pivot Moment (N-m) Mpad Pad Mass (kg-m2) n Axial Pressure Distribution Factor p Pressure in Lubricant Film (Pa) p Dimensionless Pressure ( pRC p 22 R Bearing Diameter (m) t Time (s) t Dimensionless Time ( t ) u,uu, Displacement, Velocity and Acceleration Vectors zyx ,, Direction in Horizontal, Vertical and Axial Direction ii yx , Shaft Node Displacement (m) X Shaft Center in Bearing Displacement (m) X Dimensionless Displacement of Shaft Center ( pCX / ) X Shaft Center in Bearing Velocity (m/s) X Dimensionless Velocity of Shaft Center ( pCX / ) Y Shaft Center in Bearing Displacement (m) Y Dimensionless Displacement of Shaft Center ( pCY / ) Y Shaft Center in Bearing Velocity (m/s) Y Dimensionless Velocity of Shaft Center ( pCY / ) z Dimensionless Axial Direction ( 2Lz ) Bearing Pivot Displacement (m) Dimensionless Pivot Displacement ( pC/ ) Bearing Pivot Velocity (m/s) Dimensionless Pivot Velocity ( pC / ) Bearing Pad Displacement (rad) Dimensionless Pad Displacement ( pCR / ) Bearing Pad Velocity (rad/s) Dimensionless Pad Velocity ( pCR / ) Bearing Eccentricity Local Rotational Coordinate \u03b8xi, \u03b8yi Shaft Node Rotational Displacement (rad) \u03b8p Pivot Location (rad) \u03b81, \u03b82 Pad Starting and End Angle (rad) \u00b5 Lubricant Dynamic Viscosity (Pa-s) \u03c9 Shaft Rotational Speed Inside Bearing (rad/s) Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/16/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2013 by ASME \u03a9 Rotational Speed of Rotor (rad/s) REYNOLDS EQUATION The Reynolds equation represents the main governing equation to describe the behavior of the tilting-pad bearing. The cross section of a tilting-pad bearing geometry in fixed global Cartesian coordinate with its origin at the bearing center is shown in Fig. 2 (only one pad is shown here for illustration purposes). The z axis is perpendicular to the lane of the paper following the positive right-hand rule and represents the axial direction of the tilting-pad bearing. Subject to the thin film lubrication approximation, the governing equation of pressure profile in a tilting-pad bearing is Reynolds equation on each pad. Under the assumption of an impressible Newtonian fluid and laminar flow of lubricant, the dimensionless form of Reynolds equation is t hh z ph zL Dph 2 1 1212 323 (1) where ttpRCpChhLzzRx pp ,,,2, 22 The pressure boundary conditions in the circumferential direction and axial direction are 0,, 21 zpzp (2) 0,02/1, pzp (3) In this paper, the assumed pressure distribution profile in the axial direction is of the form [13] nzpzp 1, (4) Where n has different values for different L/D ratios", + " This dimensionless axial pressure profile is placed in Eq.4 is substituted into Eq. (1), and integrated twice in the axial direction with boundary conditions Eq. (2, 3). Then the 2nd order partial differential Reynolds equation, Eq. (1) can be rewritten as the following ordinary differential equation [15] 0 2 13 1212 2 233 1 t hhBp L Dh d pdhB d d (5) where n nB n nB 1; 3 4 21 (6) FILM THICKNESS EQUATION Consider the tilting pad bearing geometry as shown in Fig. 2. The film thickness can be described as a harmonic function. The main variables in the film thickness equation are the radial clearance of the bearing (Cb), the radial clearance of each pad (Cp), the current position (X and Y) of the shaft inside the bearing, the current pad tilt angle (\u03b4) and the current pivot position (\u03b3). In dimensionless form, the film thickness equation is: ppm YXh sincos sincos1 (7) where pppbpp CRCCCmCYYCXX /,/,/1,/,/ It is important to mention that the pressure solution form, Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001441_978-3-319-05203-8_76-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001441_978-3-319-05203-8_76-Figure5-1.png", + "caption": "Fig. 5 a Thin walls, b rectangular holes and c cylindrical holes identification", + "texts": [ + " The measured surface roughness of the test parts can be divided into three groups. Surfaces with angle 10 \u201390 , where the surface roughness did not significantly change. Surfaces with angle 100 \u2013130 , that lies on the material and surface roughness increases with its increasing inclination. And surfaces, which surface roughness cannot be measured or is too high, those are with angle 140 \u2013 180 . Due to excessive roughness and the necessity of subsequent machining it is suitable to use supports at all surfaces with angle of 140 \u2013180 . For internal geometrical features (Fig. 5), holes with squared and circular crosssection and thin walls, a high geometrical precision was evaluated. Inner features were not significantly affected by the calibration of the machine or by the location of the test part at the platform. In terms of precision of test part\u2019s outer size, further calibration needs to be done, to compensate the laser focus evenly over the platform. To achieve the smallest deviation of the resulting part, it is suitable to modify the parameters of the laser power and scan speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002983_b978-0-12-818546-9.00016-6-Figure11.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002983_b978-0-12-818546-9.00016-6-Figure11.3-1.png", + "caption": "Figure 11.3 Manual wheelchairs [68]: (A) dependent; (B) independent. Source: From https://www.publicdomainpictures. net/en/view-image.php?image5 167689&picture5wheelchair; and https://www.publicdomainpictures.net/en/view-image.php? image5 167675&picture5wheelchair.", + "texts": [ + " Among wheeled mobility devices, wheelchairs, including passive and active, are recorded the most commonly used assistive devices. Wheelchairs provide the ability for patients suffering severe mobility disorders to sit and transport indoor and outdoor. Users\u2019 functional ability, device features, and environmental constraints such as the size of doorways and type of floor cover should be taken into account in selecting the wheelchair. Mainly, there are two main types of manual wheelchairs: dependent and independent. Dependent-propulsion type of manual wheelchairs (Fig. 11.3A) is mostly used for the patients who need third-party assistance in embarking, disembarking, in addition to propelling the device. They are mostly used in healthcare centers. Independent-propulsion type of manual wheelchairs refers to the wheelchairs that are propelled by the patient without any third-party assistance (Fig. 11.3B). These wheelchairs are presented in two types of frames: foldable and rigid. The frame is usually made of aluminum, titanium, manganese, and graphite that weighed half of steel-made wheelchairs [69]. In comparison between lightweight and standard wheelchairs, the prior one presented a reduction in the frequency of vibrations in parallel with increasing frequency of casters\u2019 floor detaching [70]. Lightweight wheelchairs can reduce the pain in the upper extremities effectively in users having spinal cord problems [71]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003607_tim.2020.3029365-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003607_tim.2020.3029365-Figure2-1.png", + "caption": "Fig. 2. Experimental setup for data collection from a ZD6-type railway switch system.", + "texts": [ + " To evaluate the proposed methodology, a group of switch systems in the Xiliang section of Guangzhou Metro in China were selected. An embedded data acquisition system was designed, aiming to minimize the disruptions to normal railway operation. Details of the experiment are introduced in the following subsections. The point machines that are operating in Xiliang section of Guangzhou Metro belong to ZD6-type (DC powered point machine). A schematic diagram of this type of point machine is depicted in Fig. 2(a). A load pin was applied to measure the tension and compression forces in the drive rod. An electrical current and two high internal resistance voltage transducers were fitted in the motor box to measure DC current and voltage levels, respectively. There are multiple failure modes in a railway switch system. In this paper, a drive rod overdriving fault with varying severities is considered, which is responsible for around 9% of switch failures across Great Britain [28]. Normally, the periodic operation of a railway switch is divided into four phases: inrush, unlock, move, and lock", + " Overdriving is a condition that happens in the lock phase when the force between the switch rail and the stock rail is above the ideal range. As such, a greater electric power and force can be observed when locking the switch rails at the current position due to the overdriven switches; this type of fault is used in the case study in Section III. B. The simulated overdriving fault conditions with incremental severities were achieved by fastening the hexagonal screw nut mounted on the drive rod, marked as \u2018drive rod nuts\u2019 in Fig. 2(a). Specifically, there are 6 faces on one screw nut. The nuts were turned one side (one face) at a time to emulate one level of overdriving fault severity. Fig. 2(b) shows photos of the experiments carried out on a ZD6 point machine at one of the testing sites. The key components and installation points of sensors are marked in accordance with the labels shown in Fig. 2(a). All the sensor data were collected and transmitted to a remote PC with a specified sampling rate of 20kHz. 160 data sets each composed of electrical current, voltage and force signals were obtained during the point machine normal-to-reverse movement operations. There are eight simulated overdriving fault severities, which are \u2018Fault free\u2019, \u2018Overdriving 3 faces\u2019, \u2018Overdriving 6 faces\u2019, \u2018Overdriving 7 faces\u2019, \u2018Overdriving 8 faces\u2019, \u2018Overdriving 9 faces\u2019, \u2018Overdriving 10 faces\u2019 and \u2018Overdriving 11 faces\u2019" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002298_101686616x14555429843762-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002298_101686616x14555429843762-Figure5-1.png", + "caption": "Fig. 5: Top\u2013bottom temperature differential effect on bridge cross-section.", + "texts": [ + "3), that the girders\u2019 rotation caused by temperature differences between the upper and lower parts of the bridge should be taken into account regarding the bearing, obviously due to the inevitable presence of a transverse truss at the support position. Applying this requirement to a typical double girder bridge, one can obtain the twisting rotation angle \u03c9 of each of the girders as: \u03c9= \u0394S 2h = \u03b1\u00d7S\u00d7\u0394Td1,2 2h \u00f012\u00de where \u03b1 is the thermal expansion coefficient of steel, S the distance between the girders, h the girder depth (see Fig. 5), \u0394S the temperature-induced variation of S between the top and bottom positions of the girder due to \u0394Td1,2. \u0394Td1,2 = \u03b3T \u00d7\u0394Tk1,2 \u00f013\u00de where \u03b3T is 1.35 and \u0394Tk1,2 is the characteristic value of temperature difference between the upper and lower parts of the bridge that may reach a value up to 20 C (\u0394Tk1 = 15, \u0394Tk2 = 5 as per Ref. [8] Table 3). The factor of 2 in the denominator of Eq. (12) is due to the thermal expansion of the transverse truss that is assumed to develop symmetrically with respect to the vertical middle axis of the bridge cross-section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003112_0954407020916991-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003112_0954407020916991-Figure8-1.png", + "caption": "Figure 8. The result of the finite-element analysis.", + "texts": [ + " So a layer of SURF 152 thermal effect units need to be set on the working surface to play a role as heat flux density. The final finiteelement model of the single sun gear tooth is shown in Figure 4. The experimental setup of the coupling system. Table 3. The temperature of some points in thermal network method and experimentation. Point 1 Point 2 Point 3 Point 4 Point 5 Temperature in thermal network method ( C) 45.9 39.5 38.0 45.7 52.4 Temperature in experimentation ( C) 46.6 34.9 35.7 51.2 51.7 Figure 6. The model of the single sun gear tooth. Figure 7, and the result of the FEA is shown in Figure 8. As shown in Figure 8, the temperature of the working surface is obviously higher than the temperature of other parts. The temperature distribution from the middle of the tooth surface to the end face is graded distribution, and it is completely symmetrical throughout the tooth width. These temperature distributions are in line with the actual situation. The temperature field distribution on the tooth profile line right in the middle of the meshing tooth surface was extracted from ANSYS, and the data were imported into MATLAB for analysis, the results are shown in Figure 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001875_icrom.2014.6990949-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001875_icrom.2014.6990949-Figure1-1.png", + "caption": "Fig. 1. The 3RPS parallel robot", + "texts": [ + " Also, this robot has been widely used in electronic gaming machines, virtual reality movies, simulations of aircraft and vehicles, and so on [10]. The paper is organized as follows. Kinematic modeling is carried out in section II. Dynamic formulation of the robot is presented in section III. Section IV explains the control of model-based impedance and its features. Finally, the simulations are presented in section V. II. KINEMATICS FORMULATION 3RPS mechanism consists of a moving and a fixed platform, which are connected to each other by three similar limbs. The structure is shown in Fig. 1. Each limb consists of a prismatic joint, following a spherical joint at the top and a revolute joint connected to the base. The structure comes up with one translational and two rotational degrees of freedom, and is driven by three linear actuators. Kinematics of the 3RPS parallel robot was extensively presented in [11]. It is assumed that the desired position and orientation of the output platform of the robot are given for inverse kinematics analysis. Input joint variables are to be found in order to achieve the desired trajectory" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000528_1.4031386-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000528_1.4031386-Figure2-1.png", + "caption": "Fig. 2 Schematic of the segmented brush seal orientations", + "texts": [ + " Manuscript received June 25, 2015; final manuscript received July 17, 2015; published online September 22, 2015. Editor: David Wisler. Journal of Engineering for Gas Turbines and Power MARCH 2016, Vol. 138 / 032501-1 Copyright VC 2016 by ASME Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/27/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use to triangular-shaped voids in the bristle pack of the segmented (SEG) brush-labyrinth sealing configuration SSB SEG (see Fig. 2). Some results presented in this work distinguish between the horizontal and vertical orientations of the cutting plane. In the case of the horizontal configuration SSB SH, the rotor eccentricity vector is in the cutting plane, while in the second case of the vertical configuration SSB SV the rotor eccentricity vector is perpendicular to the cutting plane. The no-whirl test rig and dynamic test rig at Technische Universitaet Muenchen were used for testing the sealing configuration with the segmented brush seal (see Ref", + " A straightforward way for incorporating the cutouts in the bristle pack is to make corresponding changes in the geometrical model directly. However, this method can lead to additional problems during generation of structural grids and is time consuming. An alternative, more convenient, approach used in this work describes the cutouts analytically by defining two conditional expressions. In the case of the horizontal segmentation, these expressions for the cutout at the larger clearance side (kcut1) and the cutout at the smaller clearance side (kcut2) are (see Fig. 2) kcut1 \u00bc 0; if y < 0\u00f0 \u00de y Do 2 < z < D 2 (6) kcut2 \u00bc 0; if y > 0\u00f0 \u00de D 2 < z < y\u00fe Do 2 (7) where y and z are coordinates of cell nodes in the porous medium region. Otherwise, the coefficients kcut1 and kcut2 are set to unity. By multiplying the resistance coefficients in Eq. (4) by the cutout coefficients kcut1 and kcut2, the porous medium model can be turned off in the CFD model for the specified regions in the bristle pack. The diagram on the right in Fig. 6 shows a fragment of the bristle pack where the white area indicates the region with deactivated porous medium model", + " \u00bc c0Xcr DLs p0 p1\u00f0 \u00de C (10) where Xcr \u00bc 1000 rad/s is characteristic frequency selected as a scaling factor. For given shaft rotational speed and preswirl velocity, the sealing configurations with segmented brush seals were tested under three pressure ratios (2, 4, 6) and three eccentricity ratios (0.2, 0.4, 0.6). Figure 8 shows the local aerodynamic stiffness coefficients versus the pressure ratio for the rotational speed of 6000 rpm and preswirl velocity of 150 m/s. Notations SV and SH correspond to the orientation of the cutting plane in respect to the shaft eccentricity vector (see Fig. 2). Both, direct and cross-coupled, local aerodynamic stiffness coefficients increase in absolute value as the pressure ratio increases. Decrease in the nondimensional values of stiffness coefficients is due to the applied dimensionless representation. The local direct stiffness is positive and quite small for the sealing configuration with nonsegmented brush seal. Segmentation of the bristle pack leads to the change of sign of the local direct stiffness. The sealing configurations with segmented brush seals behave more like the labyrinth seal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002246_978-3-319-15684-2-Figure1.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002246_978-3-319-15684-2-Figure1.2-1.png", + "caption": "Fig. 1.2 Tool set for axial rotary forging of hollow shaft flanges: 1, 2 forging roll, 3 mandrel, 4 shank, 5 inner mandrel, 6 stop, 7 top, 8 die, 9 cup, 10 base", + "texts": [ + " Then the blank is pressed to the die and this provides a reliable transmission of torque. During the rotary forging, the blank is shaped in the space that is formed by the die, mandrel and deforming tool\u2014forging roll. The tool set (die and forging roll) is manufactured from die steel with heat treatment for hardness HRC = 56\u201363 [6]. 2 L.B. Aksenov and S.N. Kunkin This experimental work investigated the possibility of manufacturing ax symmetric hollow parts with outer flange. A tool set for rotary forging (Fig. 1.2) was designed and manufactured for a horizontal type of mill [7]. Rotary forging was realized with the following main process parameters: \u2022 angle of inclination of the conical roll\u201410\u00b0; \u2022 speed of rotation of a spindle of the machine\u2014130 rev/min; \u2022 feed of forging roll\u20140.3 mm/rev in the beginning of the process and about 0.05 mm/rev at the end of the forming; \u2022 lubricant\u2014machine oil; \u2022 duration of rotary forging (forming time)\u201420\u201325s. Pre-machined pieces of pipe were used as blanks for rotary forging" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001423_aicera.2014.6908171-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001423_aicera.2014.6908171-Figure5-1.png", + "caption": "Fig. 5. Comparison of magnitude of Incal,Inf,InBV for differ", + "texts": [ + "004 - iCMMD) TION RESULTS induction motor is modeled in meters: 1HP, 415V, 1480Rpm, resistance R1= 12.35\u2126, Stator otor resistance (R2) = 6.991\u2126, 71H, Magnetizinginductance = tia (J) = 0.010Kgm2. The (2-s) negative sequence equivalent Simulation is done for different nce Factor. According to l Commission (IEC), Voltage en by (2) Magnitude and Angle) are the s. VUF is varied from 0% to 5% ulated in Table I.From the table tude of Incalincreases as voltage lated with balanced voltage is agnitudes of Incal, Inf and InBV es of VUF in Fig.5. I. TOR SIMULATED UNDER DIFFERENT LANCE FACTOR n(V) Incal (A) Inuv (A) Inf (A) .507 0.018 0.0148 0.018 96.07 -1.402 -2.535 -0.563 .022 0.046 0.0590 0.0155 165.6 2.37 5.533 -0.106 .357 0.089 0.098 0.0161 171.4 2.284 2.432 2.508 .977 0.142 0.145 0.019 174.3 2.251 2.382 0.597 .603 0.195 0.193 0.023 175.7 2.235 2.356 0.828 .232 0.248 0.24 0.028 176.6 2.227 2.341 0.929 .862 0.3012 0.288 0.0346 177.1 2.221 2.33 1.082 11.49 0.354 0.336 0.0405 177.6 2.217 2.323 1.155 13.12 0.407 0.384 0.0465 177.9 2.214 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001240_aim.2014.6878315-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001240_aim.2014.6878315-Figure2-1.png", + "caption": "Fig. 2. Magnetic axes of a PMSM", + "texts": [ + " Afterwards, the resonance demodulation technique is applied on the stator current spectra using the evaluated natural frequencies of the drivetrain to extract spectral components with rich fault information for the gearbox which is further processed to provide the fault indicator. Finally, the performance of the proposed technique to detect the gear faults in a planetary gearbox using MCSA is studied through simulation. A three phase PMSM comprises of (i) a fixed stator with three phase windings a, b and c, and (ii) a rotor with permanent magnet poles (Fig. 2). When a three phase AC power supply is applied to the stator windings, it produces a rotating magnetic field that interacts with the permanent magnet poles and pushes the rotor forwards. Park\u2019s transmission to describe PMSM dynamics has been described in [13]. It can be rearranged in a state-space equation representation in the rotor reference frame (d-q frame) as ( ) B\u039bA\u039b \u2032 + + \u2212 \u2212\u2212 = dsfd q d q dsr rqs d q LRv v LR LR r \u03bb\u03bb \u03bb \u03c9 \u03c9 \u03bb \u03bb \u03c9 0 & & & (1) where \u039b is the magnetic flux linkage vector with \u03bbq and \u03bbd as the corresponding magnetic flux linkages" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003161_j.addma.2020.101345-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003161_j.addma.2020.101345-Figure9-1.png", + "caption": "Fig. 9. Numerical domain for CFD calculations of the axially fed sample channels.", + "texts": [], + "surrounding_texts": [ + "In order to account for the inlet pressure drop (\u0394pin) in the experiment, CFD calculations have been performed. The commercial solver StarCCM+ [14] was used throughout the work. The calculations have been performed by employing a coupled implicit RANS solver. The chosen turbulence model was K-\u03b5-model. The discretization was done using polyhedral cells and prism layers at the boundary layers. The CFD does not account for the roughness of the channels. The roughness will lead to a transition from laminar to turbulent flow at lower Reynolds numbers and a shorter initial stretch to a fully developed flow field. However, the Reynolds numbers in the experiments are mostly above 2300. For these cases the entrance effects are mostly governed by the change in flow surface area. Additionally, the CFD solutions showed that the initial stretch is much shorter (L/D < 5) than for straight channels due to the geometry of the wavy channels. The impact of the roughness on the inlet pressure loss is therefore marginal which justifies the use of smooth walls in the CFD calculation." + ] + }, + { + "image_filename": "designv11_34_0000948_2013.41059-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000948_2013.41059-Figure1-1.png", + "caption": "FIG. 1 Tread pa t t e rn of tractor tire tested. (Dimensions noted are in millime ter*^", + "texts": [ + " This paper will deal with the tread movements under the contact area of radial and conventional tractor tires, which explains why the radial ply tires present less slippage at equal traction and less rolling resistance. A compari son was made between a radial and a conventional tire, with the same tread pattern, by measuring their movements on a smooth surface. In re gard to the conventional tire, the re sults were also interpreted and com pared with a special calculation me thod which had previously been elab orated and checked. The test was carried out by measur ing the movement of the points of the tread shown in Fig. 1 by means of high-precision optical equipment. The movement was applied to the contact area keeping load and inflation pres sure rigorously constant. The results are shown on the attach ed diagrams (Figs. 2 to 6) in which the abscissae indicate the position un- Presented as Paper No. 60-104 at the Annual Meeting of the American Society of Agricultural Engineers at Columbus, Ohio, June 1960, on a program arranged by the Power and Machinery Division. The authors\u2014ARRIGO CEGNAR and FUL CIERI FAUSTI\u2014are associated, respectively, with the tire development department and the tractor tire development department of Pirelli and associated companies, New York, N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure4-1.png", + "caption": "Fig. 4. Schematic view of stroke motion.", + "texts": [ + " Since FrontFlow/blue, ADVENTURE_Solid and ADVENTURE_Coupler have been developed to aim at handling large-scale problems, our analysis system can be also applied to them. Recently, we have been working on solving over one giga-degrees of freedom (DOFs) FSI problem using K computer [46]. In the future work, we will extend the present analysis to more realistic one which has a complex geometry. Although the extension engenders the increase of DOFs, it is straightforward to apply the analysis system. A flapping-wing motion can be described as a combination of three rotational degrees of freedom, namely stroke (yaw) (Fig. 4), pitch (feathering) (Fig. 5), and lead-lag (roll) (Fig. 6). According to one study [1], pitch plays a large part in the enhancement of lift generation. The pitch can be actively generated (active pitch) using actuators or passively generated (passive pitch) by the stroke if a deformable wing is adopted, as shown in Fig. 7. From a viewpoint of simplifying MAVs design, the passive pitch is very advantageous because of keeping the number of actuators low. The passive pitch has been investigated using experiments and numerical studies [15,17]", + " The endpoint of the leading edge of a wing, i.e., (0,0,0), is thus fixed. In the present study, stroke and lead-lag motions are actively generated. A pitch motion is generated either actively or passively. An actively generated motion is prescribed by the Dirichlet boundary condition on the leading edge \u0393l. We use sinusoidal waves to describe angular motions. A stroke motion is divided into two phases, namely an upstroke and a downstroke. The axis of strokes is the y-axis and two translation phases occur on the xz-plane, as shown in Fig. 4. The stroke angle \ud835\udf03s is defined by the following equation: \ud835\udf03s = CA\u0398s sin(2\ud835\udf0bfst) (deg), (22) where \u0398s is the maximum stroke angle, fs is the stroke frequency, t is the time, and CA is the acceleration coefficient given by the following Table 1 Material properties of rigid and deformable wings. Density (kg/m3) Young\u2019s modulus (Pa) Poisson\u2019s ratio Upper part (rigid wing) 1.0 \u00d7 103 1.0 \u00d7 1011 0.3 Upper part (deformable wing) 1.0 \u00d7 103 Eupper 0.49 Lower part 1.0 \u00d7 103 1.0 \u00d7 1011 0.3 Table 2 Material properties of surrounding fluid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure10-1.png", + "caption": "Figure 10. Condition of the posture of the robot.", + "texts": [ + " To define the condition, we reflect the posture of each unit in the modeled robot, as shown Fig. 8. Then, Pui is the coordinate of the center of the unit, and Pri is the coordinate of the roller B. Pui and Pri are given by (5) and (6), respectively. (5) (6) Here, l is the length between the center of a unit and its tip, and lr is the length between the center of the roller B and the tip of a unit. When the robot is fixed, the vector normal to the unit N, which passes through Pui, crosses the center of the sphere, as shown Fig. 10 (a). Then, Pui satisfies (7). (7) When the robot is moving, both of the rollers of the magnetic adhesion mechanism cannot adsorb mechanistically onto spherical surfaces. Therefore, we assume that roller B (Fig. 9 (a)) constantly adsorbs onto surfaces. When the robot is moving, there is a certain distance between roller B and the sphere, as shown in Fig. 10 (b). Then, Pri satisfies (8). (8) Here, R is the radius of the sphere, and r is the radius of the roller. The motions of omnidirectional locomotion on spherical surfaces are generated from the above robot model. The locomotion strategy is schematized in Fig. 11. In this strategy, we define an imaginary wave propagation line. In previous research [12], we defined the propagation line as a straight line, because we expected the robot to move only on flat surfaces. In this research, we redefine it as an arc line for the motion of the robot on spherical surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001696_978-3-319-20463-5_22-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001696_978-3-319-20463-5_22-Figure2-1.png", + "caption": "Fig. 2 View of six used pinions. a G. b RC. c CTW. d CTL. e MT. f GSW", + "texts": [ + " The test bed Fig. 3 used in this study consists of two rotating shafts, on which are mounted a pinion and a spur gear offering a gear ratio of 25/56. To compare the effectiveness of methods of analysis, we used six pinions with different fault states. The first one is referred as Good (G), whereas the others have several different types of defects: a Root Crack (RC), a Chipped Tooth in Width (CTW), a Chipped Tooth in Length (CTL), a Missing Tooth (MT), and General Surface Wear (GSW) as shown in Fig. 2. Three pinions are simultaneously mounted on the input shaft of the gearbox, the engagement change is done by a simple axial movement of the wheel on its axis Fig. 3b. The input shaft is driven by an electric DC motor controlled in rotational speed. The engine ensures a maximum speed of 3600 rpm. The output shaft is connected to a magnetic powder brake capable of generating different resistive torques. To record vibration signals, two accelerometers (sensitivity: 100 mV/g) are mounted radially, one vertically and the other horizontally on the outer surface of the bearing case of the output shaft of the gearbox as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003430_lra.2020.3015463-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003430_lra.2020.3015463-Figure4-1.png", + "caption": "Fig. 4. Relationship between main-track\u2019s motion and wall reaction force. The robot can climb up the spiral stair if motion is commanded as pattern 1 or 3. In pattern 2, the robot couldn\u2019t climb up because it couldn\u2019t obtain sufficient wall reaction force for rotation. Hence, more detailed verification is conducted for pattern 1 and 3.", + "texts": [ + " The method described herein is also applicable to counter-clockwise climbing spiral stairs. The tracked vehicle can control the forwarding and turning motions by adjusting the left and right velocities of the main tracks. When the robot climbs up the clockwise spiral stair, it should generate a forward motion and a right turn motion. Using the wall reaction force, the rotational movement in the right direction can be increased. To determine the operation of the main tracks, the authors tested three conditions of the motion command as shown in Fig. 4: 1. The same velocity for the left and right tracks (forward); 2. the velocities of both side tracks are in the same direction, but the left side track is larger (right turn left turn under the conditions of); 3. the velocities of both side tracks are in the same direction, but the right side track is larger (left turn under the conditions of). The test was conducted on the spiral stairs. Consequently, in condition 2 (under the condition of right turn), the robot could not climb the spiral stairs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002523_msf.879.1008-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002523_msf.879.1008-Figure1-1.png", + "caption": "Fig. 1 (a) Specimen orientations during processing; (b) Tensile specimens (dimensions in mm).", + "texts": [ + " No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (#71537728, University of Idaho, Moscow, USA-22/11/16,15:39:28) Processing. . Two types of specimens were built from IN625 alloy powder using an EOSINT M 290 400W Ytterbium fiber laser system (EOS GmbH, Munich Germany): 10 x 10 x 10 mm 3 specimens for microstructure evaluation and 85 x 18 x 3 mm 3 rectangular prismatic coupons with four building orientations (ZX, XY, XZ and 45\u1d52) for mechanical testing (Fig.1a). Next, the building platform with specimens was subjected to stress relief annealing (SR) at 870 \u1d52C (1h). Then, all the specimens were cut from the platform, and the rectangular prismatic coupons were additionally machined to obtain dumbbell-shaped tensile testing specimens as shown in Fig.1b. For reference purposes, some specimens were reserved after SR annealing (as-built state). The remaining specimens were divided into three groups: annealed at 980 \u1d52C (1h), annealed at 1040 \u1d52C (1h), and water-quenched; and HIPprocessed at 1120 \u1d52C (100 MPa, 4h) with natural cooling. All these treatments were conducted under argon atmosphere. It was hypothesized that low solution annealing HT should reduce the anisotropy of the as-built alloy\u2019s mechanical properties, whereas HIP should additionally strengthen the material by eliminating the microporosity generated during processing [16, 17]. Microstructure. For microstructural analysis, vertical ZX and horizontal XY surfaces (Fig.1a) of the cubic coupons were mechanically polished (1 \u00b5m grit size), electro-etched for 1-4 min under 2- 5V potential in a 70 mL HPO3 + 30 mL water solution. The metallographic analysis was conducted using an Olympus 3D laser microscope. The X-ray diffraction analysis was performed using a PANanalytical X\u2019Pert Pro diffractometer with CuK\u03b1 radiation. Mechanical properties. The dumbbell-shaped specimens with a 2 x 4 mm 2 gage section were tested in tension under a constant strain rate of 10 -3 s -1 at room temperature using an MTS 810 system coupled with an MTS 634.12 contact extensometer. (Fig.1b). Three specimens were used for each building orientation. The anisotropy of the mechanical properties was calculated as follows: max min max ( ) 100% x x x x \u2212 \u2206 = \u22c5 where x is the yield stress (YS), ultimate tensile strength (UTS) or elongation (\u03b4). Microstructure of the as-built alloy. Different aspects of the as-built alloy microstructure are illustrated in Figure 2. Fig. 2a shows a cross-hatched mesostructure stemming from the EOS laser scanning strategy involving a 67 o hatch rotation between two successive building layers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002551_fuzz-ieee.2016.7737924-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002551_fuzz-ieee.2016.7737924-Figure1-1.png", + "caption": "Fig. 1 DC controlling an inverted pendulum via a gear train", + "texts": [ + " Step 2: From (35) and (36), iU and iV are obtained. Step 3: By resolving (32), we obtain P , Q and ,iK then iK and Y are obtained from (41) and (42). Step 4: Based on (28) and (20), we obtain iF and iM , respectively. Step 5: From (29)-(31), , and i i iN L G are yielded. Step 6: Construct the observer (7) and take summation of \u02c6 ( )i t\u03b7 to obtain \u02c6( )x t . IV. ILLUSTRATIVE EXAMPLE Consider an application for DC controlling an inverted pendulum via a gear train [35], [36]. The system is shown in Fig. 1 and the armature-controlled DC motor is depicted in Fig.2. The nonlinear model of above system are described as follows 2 1 2 1 32 3 2 3 0 sin( ) 0 1 m b a a a a x x NKgx x x u l ml x K N R x x LL L = + + \u2212 \u2212 (43) where the vector of state variables is defined as [ ]1 2 3( ) , , , , TT p p ax t x x x I\u03b8 \u03b8= = , mK is the motor torque constant, bK is the back EMF constant and N is the gear ratio. The parameters of system are 29.8 /g m s= , 1 l m= , 1 m kg= , 10N = , 0.1 m/AmK N= , 0.1 /bK Vs rad= , 1 aR = \u03a9 and 100 HaL m= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003073_iet-epa.2019.0941-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003073_iet-epa.2019.0941-Figure3-1.png", + "caption": "Fig. 3 Structure of the proposed WT-IPMSM", + "texts": [ + " However, the region IV is the high-speed operation region with four parallel branches. Considering reduction of output load torque and increase of motor loss density during high-speed highfrequency condition, the excitation magnetic field must be deeply weakened. Therefore, the electrical load accounts for a larger proportion in the electromagnetic load distribution. In this paper, a WT-IPMSM with 45 kW and maximum speed of 12,500 rpm is designed for EV application. The proposed WTIPMSM shown in Fig. 3 is analysed by the finite-element method. By the WT strategy, the speed range of the proposed WT-IPMSM is divided as follows: region I (0\u20132000 rpm), region II (2000\u20134000 rpm), region III (4000\u20138000 rpm), and region IV (8000\u201312,500 rpm). The parameters of the proposed WT-IPMSM and conventional IPMSM are listed in Table 1. Obviously, the proposed WT-IPMSM has the same effective material utilisation rate as the conventional IPMSM. The difference is that the number of parallel branches (a = 1, 2, and 4) of WT-IPMSM will change with the load characteristics of each operating interval", + " Since the iron loss is affected by problems such as machining and mechanical assembly, it is necessary to add certain correction factor, shown as follows: Ph = \u2211 k = 1 \u221e kh(Bk) f k(1 + \u03b3Dk) (12) Pe = \u2211 k = 1 \u221e ke(Bk, f k) f k 2(1 + \u03b3Dk) (13) where \u03b3 is the coefficient based on iron core manufacture and Dk is the ratio of short axis to long axis of the trajectory ellipse magnetic density. The core loss analysis of 45 kW IPMSM, which coupled finiteelement analysis with analytical method has been done in subregional operation. The main magnetic field in the tooth of stator core is alternating magnetisation while field in the yoke is rotational magnetisation. Therefore, the tooth and yoke are divided into four regions to calculate hysteresis and eddy current loss, respectively, as shown in Fig. 3. The corresponding results are shown in Fig. 9. From Fig. 9, the distribution of stator core loss for two different types (WT-IPMSM and IPMSM) are same in region I and region II, IET Electr. Power Appl., 2020, Vol. 14 Iss. 7, pp. 1186-1195 \u00a9 The Institution of Engineering and Technology 2020 1189 Authorized licensed use limited to: University of Massachusetts Amherst. Downloaded on July 05,2020 at 09:26:27 UTC from IEEE Xplore. Restrictions apply. respectively. Since both WT-IPMSM and IPMSM adopt MTPA control strategy, the excitation magnetic field can produce constant output torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000395_s10409-015-0399-4-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000395_s10409-015-0399-4-Figure7-1.png", + "caption": "Fig. 7 Three modes of response of hinged rod. a Mode I. b Mode II. c Model III", + "texts": [ + " Figure 5a, 5b shows the effects of these three parameters on r . Obviously, r is independent of k1, and usually r > 1 except for small regions where k2 is very small. The maximum value rmax = 1.509 occurs at k2 = 0.25 and \u03b20 = 1.60 rad. For l = 0.20 m, during the response of the hinged rod, the velocity reaches its maximum of 0.5914 m/s at time t = 2.36 s, at position x = 0.014 m and y = 0.3997 m (Fig. 6). When the hinged rods with different k2 and \u03b20 are released, the responses can be classified into three modes (see Fig. 7, where the arrows indicate the directions of the angular velocities). Mode I occurs under a relatively large k2, that is, the hinged rod behaves like a rigid rod apart from a slight vibra- tion between two segments.Mode II occurs under a relatively small k2; a peak and a valley exist in the responses of \u03b2\u03071 and \u03b2\u03072, respectively. The maximum velocity ratio r is achieved by this mode. Mode III occurs if both k2 and \u03b20 are very small, which will result in the two segments\u2019 making contact with each other", + "201 occurs when \u03b20 = \u03c0 /2. The rotation of the rod is depicted in Fig. 8: the time histories of \u03b21, \u03b22, \u03b2\u03071, \u03b2\u03072 are plotted in Fig. 8a, while the responses of the x- and y-coordinates and the velocity of P2 are plotted in Fig. 8b. As k2 increases, the vibration decreases, and thus the response of the hinged rod will be closer to that of the rigid rod; that is, r will tend to 1.0. The ratio r reaches its maximum when k2 = 0.25 and \u03b20 = 1.60 rad, while the response can be classified as mode II, as shown in Fig. 7b. The angle between two segments, i.e., \u03b21\u2212\u03b22, first decreases and then increases owing to the inertial force and springG2. At a certain time before the rod becomes straight again, the rod reaches a state where 0 < \u03b21 \u2212 \u03b22 < \u03c0 /2, \u03b2\u03071 < 0, and \u03b2\u03072 > 0, then the absolute values of \u03b2\u03071 and \u03b2\u03072 rise rapidly, as shown in Fig. 9. This is because in this situation, the dominating terms in Eq. (4) are \u03b2\u03072 1 and \u03b2\u03072 2 , so Eq. (4) can be simplified to ( \u03b2\u03081 \u03b2\u03082 ) = sin (\u03b21 \u2212 \u03b22) 4/9 \u2212 1/4 cos (\u03b21 \u2212 \u03b22) 2 \u00d7 (\u2212 1 6 \u03b2\u0307 2 2 \u2212 1 4 cos (\u03b21 \u2212 \u03b22) \u03b2\u03072 1 1 4 cos (\u03b21 \u2212 \u03b22) \u03b2\u03072 2 + 4 3 \u03b2\u0307 2 1 ) ", + " (5)) is derived as V\u0307 2 l2 = 2 ( \u03b2\u03071\u03b2\u03081 + \u03b2\u03072\u03b2\u03082 ) + 2 cos (\u03b21 \u2212 \u03b22) ( \u03b2\u03081\u03b2\u03072 + \u03b2\u03071\u03b2\u03082 ) \u22122 sin (\u03b21 \u2212 \u03b22) \u03b2\u03071\u03b2\u03072 ( \u03b2\u03071 \u2212 \u03b2\u03072 ) . (8) According to Eq. (7), when \u03b21 = \u03b22, \u03b2\u03081 and \u03b2\u03082 are almost zero. Therefore, by Eq. (8), the free-end velocity will approach its maximum at this time. Because of this mech- anism, the maximum velocity of the free-end P2 becomes considerably higher owing to the existence of the middle hinge. If both k2 and \u03b20 are very small, then the two segments of the rod may come into contact with each other, as shown in Fig. 7c. In this mode, \u03b22 decreases monotonically because spring G2 cannot draw it back, as shown in Fig. 10. A map of the modes is shown in Fig. 11. The boundary betweenModes I and II is at k2 \u2248 1.1, and the value will vary slightly for different \u03b20. Mode III only exists in a very small area in the bottom left corner of the map, so its boundary is not plotted. In this section, the responses of the rod with more hinges are discussed, while the effects of other factors, such as the damping effect, the position of the hinge, and the concentrated mass at the free end, are also considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.9-1.png", + "caption": "Figure 9.9 First kind of motion singularity of a Puma manipulator.", + "texts": [ + "82) implies that the first kind of motion singularity occurs if k1 = d4s\ud835\udf0323 + b2c\ud835\udf032 = 0 (9.96) On the other hand, referring to Part (c) of Section 9.1.2, it is seen that d4s\ud835\udf0323 + b2c\ud835\udf032 = r\u20321 (9.97) Recall that r\u20321 is the projection of the vector r\u20d7SR on the first axis of the link frame 1(O). Therefore, in this singularity with r\u20321 = 0, r\u20d7SR becomes parallel to the axis of the first joint. In other words, the wrist point R becomes located on a cylindrical surface, whose radius is d2. This surface is a noticeable feature of this singularity. This singularity is illustrated in Figure 9.9. The singularity surface mentioned above also constitutes a workspace boundary for the wrist point. That is, as also indicated in Part (a) of Section 9.1.3 by Inequality (9.41), the wrist point cannot get into the region bounded by this cylindrical surface. Note that the appearance of this motion singularity becomes the same as the appearance of the first kind of position singularity if d2 = 0. If the manipulator has to pass through this singularity or keep on moving on the singularity surface as a task requirement, then, according to Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure7.8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure7.8-1.png", + "caption": "Fig. 7.8 Capacitive machine model, drawings acc. [7.1]", + "texts": [ + " The capacitor is nearly short circuited due to the high frequency. 118 7 Converter Caused Shaft Voltages Machines with a relatively large capacitive coupling between stator winding and rotor have often a remarkable capacitive coupled shaft voltage beside the circulating flux phenomenon. This is more typical for asynchronous machines than for synchronous machines. The shaft voltage level depends mainly upon the individual distribution of capacities. An overview of the dominating capacities is given in fig. 7.8. The stator winding is coupled via the air-gap to the rotor. Of course this capacity is smaller than the coupling to the stator frame. Capacities occur between the rotor and the bearings as well as between the rotor via the air-gap to the stator core in series to the stator winding rotor coupling. A sinusoidal three-phase voltage system is coupled to the rotor as well, but the rotor would be similar to a neutral point, which is created by capacities. In case of a common mode voltage the situation is different" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001343_pedstc.2013.6506688-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001343_pedstc.2013.6506688-Figure7-1.png", + "caption": "Figure 7. control diagram", + "texts": [ + " 5) the electrical speed of the outer rotor is equal to the stator voltage frequency and the electrical speed of the inner rotor is equal to the difference between the stator and inner rotor voltage frequencies. In such condition inner rotor field, outer rotor field and stator field are all rotate with the same angular speed in DMP machine volume. Moreover, as it is shown in fig. 6, the torque caused in each rotor is equal to the mechanical load torques. In the next step, a DMP is simulated in a close loop test and with a PI controller. The control algorithm is shown in fig. 7. In this test, speed references and mechanical load torques of both rotors are the inputs of the system. Fig. 8 shows the variation of the inputs. Inner rotor speed, outer rotor speed, inner rotor current and stator current are considered as the outputs of the system. Simulation results are as follows: as the simulation results show, speed of both rotors track their refrence speed in fig. 9. According to fig. 10, stator current frequency is equal to the outer rotor electrical speed. Inner rotor current frequency is also equal to difference between the outer rotor electrical speed and the inner rotor electrical speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003295_s00170-020-05693-0-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003295_s00170-020-05693-0-Figure1-1.png", + "caption": "Fig. 1 Illustration of the flank parameters of a conical end-mill", + "texts": [ + " 4, we established an optimization model and put forward the geometrical constraints to avoid interference and abnormal flank profile based on the cylindrical end-mill. Afterward, an integral procedure was built to optimize and simulate the conical flank grinding. In Sec. 5, a numerical simulation was carried out. The rationality and accuracy of the above models were verified by comparing the case results. In Sec. 6, the contribution of the work were summarized. As an important feature of a conical end-mill, conical flank grinding is often defined in the industry with two parameters, flank phase angle \u03b8f and relief angle \u03b1, which are illustrated in Fig. 1. For the conical end-mill, the radius is denoted as rT; \u03c6 is the cone angle and L is the length of the tool. The helical angle \u03bb is the angle between the tangent of cutting edges and the generatrix of the conical end-mill at the cutting edge point. The radius rT can be expressed in Eq. (1): rT \u00bc rT0 \u00fe l tan\u03c6; \u00f01\u00de where rT0 is the initial radius of the conical end-mill, l \u2208 [0, L]. The conical flank surface is not designed in particular, while it is generated by the grinding wheel during machining" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000355_pc.2015.7169988-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000355_pc.2015.7169988-Figure1-1.png", + "caption": "Fig. 1. Magnetic levitation system.", + "texts": [ + " The proposed parametric reference governor setup is applied to a fast and unstable magnetic levitation system, controlled by a low-cost Arduino microcontroller. We show that despite the microcontroller only providing modest computational resources and a small amount of memory storage (2 kilobytes), it can accomodate the parametric reference governor and execute it under hard real-time bounds. We present experimental results that confirm efficiency of the proposed solution. We consider a magnetic levitation system, depicted in Fig. 1, in which the vertical displacement of a metal ball is controlled by manipulating the voltage injected to a coil. The position is controlled by a conventional PID controller which, however, does not explicitly take position constraints into account. To enforce constraint satisfaction, we employ a reference governor setup as shown in Fig. 2. The governor shapes the desired reference r in such a way that the PID controller, which takes the regulation error based on the shaped reference w as its input, generates control commands which lead to closed-loop profiles that obey position constraints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002831_ls.1497-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002831_ls.1497-Figure4-1.png", + "caption": "FIGURE 4 Static pressure contour plot for the journal bearing (\u03b5 = 0.33 and L/D = 0.66) [Color figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " In this context, study had been made under dynamically loaded conditions to understand the wear of the bearings. So FSI technique was used to compare the simulated wear results with realistic conditions. Additionally, the focus was given on the deformation of the bearing under boundary lubrication and elastohydrodynamic lubrication conditions. The calculated L/D ratio for the journal bearing is 0.66, and an eccentricity ratio of 0.33 was assumed for initial analysis to understand the pressure distribution when the bearing is under the influence of external load of 30 kN. Figure 4A,B shows the static pressure contour plot for \u03b5 = 0.33 and L/D ratio = 0.66 with and without cavitation model approach. The simulation results indicated that maximum static pressure of 0.371 MPa without the cavitation model concept and 0.377 MPa with the cavitation model concept was observed in a region closer to the convergent portion in the bearing. In the Figure 4A, the pressure distribution is changed to pressure build-up, and the pressure drop has vanished. However, in Figure 4B, the pressure distribution showed pressure build-up and pressure drop in the bearing. The pressure build-up lies closer to the convergent portion and ends after the smallest gap that leads to a different reaction on the shaft. From the study, it is confirmed that the magnitude of the maximum pressure build-up in journal bearings is almost same for with and without cavitation models. This could be due to the chosen eccentricity value and the rotational speed of the shaft is less for the studied bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002329_med.2016.7536052-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002329_med.2016.7536052-Figure2-1.png", + "caption": "Figure 2. Mechanical Analysis of Quadrotor", + "texts": [ + " It is important to point out that the two methods are not isolated, and they can be combined to realize a better performance. The parameters of the quadrotor used in this paper are shown in table I. In this paper, dynamics model and measurement model will be given at first. Then PID controller will be introduced briefly. At last, faults in a rotor will be considered, which will influence the thrust coefficient k (unit ) and moment coefficient b (unit ). FDD and controller reconfiguration mechanism will be designed for the fault. Before giving the dynamics model, let define the coordinate system. As shown in Fig. 2, body coordinate system and inertial coordinate system are introduced, and vectors in one system can be transformed to ones in another system by multiplying them with transformation matrix. The rotation sequence from the inertial system to the body system is z, y and x. The forces concerned are the gravity of quadrotor and the lifts of the four rotors, and the moments concerned are the torques of the four rotors, while other forces and moments are too small to be considered [6], [9]. Dynamics model are given from (20) to (22)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002969_978-3-030-40172-6_19-Figure21.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002969_978-3-030-40172-6_19-Figure21.4-1.png", + "caption": "Fig. 21.4 Process model discovered by ProM\u2019s Inductive Miner. Note that all activities occur once per procurement order", + "texts": [ + " However, the three middle activities are executed in any order. For example, in rare cases the organization pays before receiving the invoice and goods. Hence, there are six different process variants. In our event log there is information about 2000 orders, i.e., in total there are 10,000 events. The most frequent variant is hplace order, receive invoice, receive goods, pay order, closei which occurs 857 times. The least frequent variant is hplace order, pay order, receive goods, receive invoice, closei which occurs only four times. Figure 21.4 shows the process model discovered by ProM\u2019s Inductive Miner [2]. The three unordered activities in the middle are preceded by an AND-split and are followed by an AND-join. The model correctly shows that all activities happen precisely 2000 times. Applying other basic process discovery algorithms like the Alpha algorithm and region-based techniques yield the same compact and correct process model (but then often directly expressed in terms of Petri nets). Let us now look at the corresponding Directly-Follows Graphs (DFG) used by most commercial Process Mining tools [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002593_978-3-319-46155-7_6-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002593_978-3-319-46155-7_6-Figure8-1.png", + "caption": "Fig. 8 Kinematics lever of a business class seat", + "texts": [ + " SLM opens opportunities to fabricate complex optimization designs without any adjustments after optimization. The reduction of weight is an important factor in aerospace industry (Rehme 2009). The weight of the aircraft determines fuel consumption. An aircraft seat manufacturer is investigating the opportunities to save weight in their business class seats through SLM. One part of the seat assembly, a kinematics lever, was selected to investigate the potential of the direct fabrication of topology optimization results via SLM (see Fig. 8). In a first step, the maximum design space and connecting interfaces to other parts in the assembly were defined (see Fig. 8) to guarantee the fit of the optimization result to the seat assembly. Interfacing regions are determined as frozen regions, which are not part of the design space for optimization. The kinematics lever is dynamically loaded if the passenger takes the sleeping position. Current topology optimization software is limited to static load cases. Therefore, the dynamic load case is simplified to five static load cases, which consider the maximum forces at different times of the dynamic seat movement. Material input for the optimization is based on an aluminum alloy (7075) which is commonly used in aerospace industry: material density: 2810 kg/m3, E Modulus: 70" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001651_j.ifacol.2015.10.261-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001651_j.ifacol.2015.10.261-Figure2-1.png", + "caption": "Fig. 2. Path-following framework.", + "texts": [ + " From a technical point of view, the proposed cooperative motion procedure is based on: (1) a Lyapunov-based virtual target path-following algorithm that allows a single vehicle to approach and follow a generic reference path in the horizontal plane; (2) a distributed coordination scheme allowing the motion synchronization of the two vehicles, using a minimum data exchange (i.e. the values of the virtual target curvilinear abscissae, as described afterwards). The design of a virtual target based guidance system is reported in Bibuli et al. [2009], a brief description of the guidance approach is given in the following. With reference to Fig. 2 and assuming the vehicle\u2019s motion restricted to the horizontal plane, the task is to reduce to zero both the position error vector d, i.e. the distance between the vehicle and the virtual target attached to the Serret-Frenet frame < v >, and the orientation error \u03b2 = \u03c8 \u2212 \u03c8p, where \u03c8 and \u03c8p are the vehicle\u2019s direction of motion and local path tangent respectively, expressed with respect to the earth-fixed reference frame < w >. Performing some geometrical computations, the following kinematic error model is obtained and expressed with respect to the frame < v >: s\u03071 = \u2212s\u0307 (1\u2212 ccy1) + U cos\u03b2 y\u03071 = \u2212ccs\u0307s1 + U sin\u03b2 \u03b2\u0307 = r \u2212 ccs\u0307 (1) where r = \u03c8\u0307, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003104_1.i010782-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003104_1.i010782-Figure2-1.png", + "caption": "Fig. 2 Different measurements from UAV i.", + "texts": [ + " The possible sensors on each UAV are the range, bearing, and speed sensors. The states to be estimated are the position xT; yT \u22a4 and velocity _xT; _yT \u22a4 of the target. The sensor models of the respective sensors are given as rT;i xT \u2212 xi 2 yT \u2212 yi 2 q (22) \u03b8T;i tan\u22121 yT \u2212 yi xT \u2212 xi (23) vT;i vT cos \u03b8T;i \u2212 \u03c8T \u2212 vi cos \u03b8T;i \u2212 \u03c8 i (24) The outputs of the range, bearing, and speed sensors are rT;i; \u03b8T;i, and vT;i, respectively. Expressions of these measurement models are given in the literature [15,17,18,21]. The variable vT;i in Fig. 2 denotes the component of relative velocity between the target T and UAV i, along the normal direction to the line of sight. We need to use the expressions of vT;i and vT;i to analyze the optimal conditions derived from the FIM. For convenience, we rewrite vT;i from Eqs. (23) and (24), and we provide vT;i as obtained from Fig. 2. These expressions are given as follows: vT;i _xT \u2212 _xi cos \u03b8T;i _yT \u2212 _yi sin \u03b8T;i (25) vT;i _xT \u2212 _xi sin \u03b8T;i \u2212 _yT \u2212 _yi cos \u03b8T;i (26) We use vT;i and vT;i, expressed in Eqs. (25) and (26), to compute the FIM. Although, vehicle i has knowledge of vT;i from the speed sensor and vT;i is unknown to i. Next, the general expression of the FIM is determined to obtain the conditions for optimal target\u2013UAV formation. The Jacobian \u2202hT;i\u2215\u2202XT \u2208 R3\u00d74 is determined for the aforementioned sensor models [Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000185_ijpt.2018.090374-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000185_ijpt.2018.090374-Figure3-1.png", + "caption": "Figure 3 Meshing cyclic power loss with different surface topography and crowning (see online version for colours)", + "texts": [ + " Several investigators have studied improvements in the meshing contact stress distribution for misaligned spur gears through crowning (Seol and Kim, 1998; Simon, 1989). In order to study the effect of crowning on power loss, the amount of crowning is varied between 50%\u2013150% of the current in-field design. The amounts of applied crowning used in the current study are listed in Table 1. In order to study the simultaneous effect of crowning and surface roughness on power loss, a map of these values is generated. Figure 3 shows the total power loss with different crowning and surface roughness. Figure 4 represents percentage change in the total power loss with respect to the current base design. Referring to Figures 3 and 4, the power loss of the planetary gear sets can be reduced by as much as 5%. However, considering the high efficiency of these gear sets, the absolute value of this reduction only amounts to 45 W per meshing cycle. This gain in efficiency should be set against the entailing manufacturing costs, indicating little incentive for implementation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002593_978-3-319-46155-7_6-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002593_978-3-319-46155-7_6-Figure7-1.png", + "caption": "Fig. 7 New multi-spot SLM system based on diode lasers", + "texts": [ + " Scanning strategy 2: the two laser beam sources and the two laser-scanning systems expose the same build area. Again, a doubling of the build-up rate is achieved. In addition, new process strategies may be developed: a laser beam is used for preheating the powder material that is followed by a second laser beam that melts the powder afterwards. Another way to increase productivity is by parallelization of the SLM process using multiple laser beam sources and multi laser-scanning systems in one machine (see Fig. 7). The authors have developed a new way, based on multi-diode lasers. The machine design uses no scanner system, and relies on a printer head instead, featuring several individually controllable diode lasers that move using linear axes. The advantage of multi-spot processing: the system\u2019s build-up rate can be increased by adding a virtually unlimited number of beam sources\u2014with no need for modifications to the system design, to the exposure control software or to process parameters. The new machine design allows to increase building space, simply by extending the travel lengths of the axis system, without changing the optical system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure6.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure6.3-1.png", + "caption": "Figure 6.3 Two versions of a universal joint.", + "texts": [ + " It actually consists of two independent revolute joints that connect three kinematic elements, which are ab, bab, and ba. Here, bab is an intermediate kinematic element between the main kinematic elements ab and ba, which belong to the members a and b that are connected by the universal joint. Because of the two independent revolute joints, the mobility of a universal joint is \ud835\udf07ab = 2. There are basically two versions of a universal joint, which are distinguished by the arrangement of the axes of the first and second revolute joints. The two versions are illustrated in Figure 6.3. For both versions, the characteristic location equation is r(ab) ab,ba = 0 (6.16) For the version shown on the left-hand side of Figure 6.3, the rotation sequence is ab rot[u\u20d7(ab) k = u\u20d7(bab) k ,\ud835\udf19ab] \u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2192bab rot[u\u20d7(bab) i = u\u20d7(ba) i ,\ud835\udf03ab] \u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2192ba (6.17) The corresponding orientation equation is C\u0302(ab,ba) = eu\u0303k\ud835\udf19ab eu\u0303i\ud835\udf03ab (6.18) For the version shown on the right-hand side of Figure 6.3, the rotation sequence is ab rot[u\u20d7(ab) i = u\u20d7(bab) i ,\ud835\udf03ab] \u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2192bab rot[u\u20d7(bab) k = u\u20d7(ba) k ,\ud835\udf19ab] \u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2192ba (6.19) The corresponding orientation equation is C\u0302(ab,ba) = eu\u0303i\ud835\udf03ab eu\u0303k\ud835\udf19ab (6.20) Joints and Their Kinematic Characteristics 131 A spherical joint is also known as a globular joint or a ball-and-socket joint. Indeed, it consists of two kinematic elements that have the shapes of a ball and a socket as illustrated in Figure 6.4. Within the encasement of the socket, the ball can rotate freely about any arbitrary direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000073_iccs45141.2019.9065331-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000073_iccs45141.2019.9065331-Figure2-1.png", + "caption": "Fig 2: - Mathematical modeling of TRMS", + "texts": [ + " In Segment IV, Fractional Order Sliding Mode Control for TRMS is elaborated where its sub-section A formulates the control law, and the sub-section B shows the Lyapunov stability and convergence in finite time. Segment V depicts the simulation results and in Segment VI, Conclusion is drawn for the above study TRMS shown in Fig. 1 is developed in Feedback Instrument Ltd., has been subject of several research work. The mathematical model of TRMS is proposed by considering the following assumptions. First, using differential equations of first order, a subsystem dynamics. Secondly, from aerodynamic stream laws system the subsystem for the propeller is designed. Fig. 2 presents TRMS\u2019 mathematical modelling. The nonlinear equations for vertical plane are:: (1) where, -nonlinear static characteristic (2) -gravity momentum (3) -fractional force momentum (4) - Gyroscopic momentum (5) The motor momentum is described by an approximated first order transfer function to Laplace Domain: (6) The nonlinear equations for horizontal plane are: (7) where, - nonlinear static characteristic (8) - friction forces momentum (9) The cross momentum is approximated by (10) 978-1-5386-8113-8/19/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002456_hnicem.2014.7016234-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002456_hnicem.2014.7016234-Figure2-1.png", + "caption": "Fig. 2. Motion analysis of the CNC router system", + "texts": [ + " The MakiBayan program aims to strengthen the research and development sector to provide solutions of industry problems in metals engineering, furniture and equipment fabrication. It also aims to work more closely and share resources to work common goals with private R&D partner. The program started with the development of the two (2) machines: the CNC router dubbed as the \u2018Super Lilok\u2019 machine and the CNC plasma cutter as \u2018Plasmanoy\u2019 that uses plasma in cutting sheet metals. For the CNC router operations, its motion can move and cut in three (3) directions (X, Y and Z) as shown in Figure 2. The X-axis is running left to right. The Y-axis runs from front to back and the Z-axis for the up and down. There is also an optional rotary axis (A-axis) that is integrated in the machine that rotates along the X-axis. Figure 3 provides an illustration of the signal flow that occurs during transformation of the drawing file to a machine code that will be read and understand by the controller [3]. Just like any CNC machines, the CNC router uses a computing software and breakout board or motion controller (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000318_j.finel.2015.07.008-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000318_j.finel.2015.07.008-Figure9-1.png", + "caption": "Fig. 9. FE model of the geared rotor system.", + "texts": [ + " X \u00bc \u00bdu1; v1;w1;\u03b8x1;\u03b8y1;\u03b8z1;u2; v2;w2;\u03b8x2;\u03b8y2;\u03b8z2;\u2026;un; vn;wn;\u03b8xn;\u03b8yn;\u03b8zn T \u00f029\u00de With the proposed 6-DOF modeling method, an unbalance response analysis is performed with a 600 me turbo-chiller rotor bearing system in Ref. [16], which is shown in Fig. 7. When a test unbalance of 19.7 a mm is attached to the impeller, the maximum coupled unbalance responses at the motor rotor through the 6-DOFs modeling method in this paper together with the results of Ref. [16] are shown in Fig. 8. The unbalance responses in this paper are consistent with the results in the reference. The FE model of the five-shaft geared rotor system is shown in Fig. 9, where 169 beam elements and 174 nodes are included and shown in Table 1. The gears, couplings, and impellers are simplified as lumped mass elements. The parameters of the gears and the bearings on the both sides are same for each shaft and shown in Tables 2 and 3 respectively. According to Eqs. (26) and (27), the radial stiffness and damping values of each bearing (in the y- and z-direction) respectively are calculated in Fig. 10, where the effect of gravity is considered and the force direction is along the y-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000149_978-3-030-15187-4_1-Figure1.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000149_978-3-030-15187-4_1-Figure1.4-1.png", + "caption": "Fig. 1.4 The relative sensings, which are expressed in local coordinate frames, between neighboring agents", + "texts": [ + " It is important to find a relationship between ziji = pij and zii j = \u2212pij where the former pij is the j th position with respect to i , expressed in i\u03a3 , and the latter \u2212pij is the i th position with respect to j , expressed also in i\u03a3 . So, pij and \u2212pij can be expressed in g\u03a3 as (pij ) g = Ri g p i j (\u2212pij ) g = Ri g(\u2212pij ) From theobservation (pij ) g = \u2212(\u2212pij ) g ,wehave Ri g p i j = \u2212Ri g(\u2212pij ),which implies that pij = \u2212(\u2212pij ). Thus, we can see that the j th position measured in i\u03a3 and expressed in i\u03a3 , and the i th position measured in j\u03a3 but expressed in i\u03a3 are the same vector with the opposite direction. Consequently, the negation of \u2212pij , i.e., \u2212(\u2212pij ), can be expressed as \u2212(\u2212pij ) = pij . Figure 1.4 depicts the relative sensings between neighboring agents i and j . Let an agent i be updated under the single-integrator dynamics as p\u0307i = ui The above equation means that the agent i is dynamically updated as per the input ui . We need to distinguish the updates in g\u03a3 and i\u03a3 . The equation p\u0307i = ui represents an update in g\u03a3 because ui is given in g\u03a3 . But, if the control input ui is SE(d) invariant in d-dimensional space [39], by transforming the orientation to local frame as we can have p\u0307ii = uii which is the update in i\u03a3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000415_3527600906.mcb.20130069-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000415_3527600906.mcb.20130069-Figure12-1.png", + "caption": "Fig. 12 Surface-generated acoustic wave (SGAW) sensor set-up. The arrows at the top indicate the flow of the liquid sample (1) in which the sensor is immersed. The elements of the SGAW biosensor are a piezoelectric crystal (2), IDTs (3), the surface acoustic wave (4), and immobilized antibodies (5)", + "texts": [ + " SAW devices operate at higher frequencies than BAW devices which, in principle, may lead to higher sensitivities because the acoustic wave penetration depth in the adjacent media is reduced [105]. In a typical configuration, an electrical signal is converted at the input IDT into a polarized transverse acoustic wave traveling parallel to the substrate surface. The amplitude and/or velocity of the wave are affected by any coupling reaction at the surface. The output IDT at the opposite end picks up the acoustic wave and converts it back to an electrical signal; any attenuation of the wave is then reflected in the output signal (Fig. 12). Depending on the piezoelectric substrate material, the crystal cut, the positioning of IDTs on the substrate, plate thickness and corresponding to the analyte molecules (6) in the sample. The driving electronics (7) operate the SAW biosensor and generate changes in the output signal (8) as the analyte binds to the sensor surface. Reproduced with permission from Ref. [107]; \u00a9 2008, Springer Science and Business Media. wave guide mechanism, different operational modes of SGAW such as shear horizontal surface acoustic wave (SH-SAW), surface transverse wave (STW), Love wave, shear horizontal acoustic plate mode (SH-APM), and layer-guided acoustic plate mode (LG-APM) can be achieved [95]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003422_aim43001.2020.9159044-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003422_aim43001.2020.9159044-Figure3-1.png", + "caption": "Fig. 3: Finger attachment design.", + "texts": [ + " The Realsense camera is used to model and track the shape of the cable, while the Kinect camera is used to estimate the socket pose and filter the point cloud. This scenario corresponds to a humanoid robot equipped with a depth camera at its left arm wrist and another depth camera at its head. One arm performing the task while the other is providing the perception feedback. In addition, a custom attachment is mounted on the fingers of the gripper to enable a formclosure grasp around the cable cross section. Figure 3 shows this attachment in SolidWorks and mounted on the fingers. From a system design perspective, in order to complete the task autonomously, the system needs to be able to model a linear flexible object, to keep track of the cable configuration and to detect the pose of the target socket through pose estimation, to plan motions to move the robot to desired configurations, and to align the cable-tip with the target socket by a controller. A flow of the proposed system architecture is presented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000442_coase.2015.7294131-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000442_coase.2015.7294131-Figure10-1.png", + "caption": "Fig. 10: A redundant quasi-planar robot having 8 revolute joints.", + "texts": [ + " The reason is the following: for non\u2013 planar manipulators a rigid transformation implies a rotation around an axis that, in general, depends on the current configuration (on the contrary, for planar manipulators the axis is always the same, orthogonal to the plane of the robot). As a consequence, convexity with respect to the decision variables is lost. The method, however, may be applied whenever the kinematic inversion problem, for a given position of the end-effector, can be reduced to a planar one. An example is given in Fig. 10: for a given position of the end-effector, the angle of the first revolute joint (vertical axis) is defined, hence a planar problem1 can be formulated. The possible adaptation of the method to other special classes of non\u2013planar manipulators is currently under investigation. [1] S. M. Lavalle, Planning algorithms. Cambridge University Press, 2006. [2] E. Rimon and D. E. Koditschek, \u201cExact robot navigation using artificial potential functions,\u201d IEEE Transactions on Robotics and Automation, vol. 8, no" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003205_s40430-020-02447-7-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003205_s40430-020-02447-7-Figure1-1.png", + "caption": "Fig. 1 Structural diagram of mechanical seal. Notes: 1\u2014rotor, 2\u2014 steady pin, 3\u2014stator, 4\u2014O-ring, 5\u2014casing, 6\u2014spring, 7\u2014housing", + "texts": [ + " What is more, further research is deserved to investigate the squeeze effect on hydrodynamic lubrication and clearance flow. The current paper aims to investigate the squeeze effects on cavitation characteristic and sealing performance of spiral groove mechanical seals. Meanwhile, a transient version of the modified Reynolds equation considering mass-conserving boundary algorithm is proposed based on FDM method. The results could be a reference for unsteady lubrication analysis, leakage control and dynamic stability in mechanical seals. 2 Theoretical model with\u00a0squeezing and\u00a0traction effect As shown in Fig.\u00a01, the fundamental elements of spiral groove mechanical seal are the rotor (rotational face) that is rigidly mounted on the shaft, while the stator (stationary face) is spring-loaded and floats axially [21]. The relative rotational motion and normal reciprocating motion between the stator and rotor could generate hydrodynamic pressure, namely the viscous shearing effect and squeeze Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:361 1 3 Page 3 of 11 361 effect. Meanwhile, this phenomenon could be further enhanced by a group of spiral grooves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.14-1.png", + "caption": "Figure 10.14 Stewart\u2013Gough platform.", + "texts": [ + "3 can be commanded to have an appropriate desired value because \ud835\udf033 is an active joint variable. Then, Eq. (10.299) gives the corresponding value of ?\u0307?6. Thus, the indefiniteness caused by MSIK-3 can also be eliminated just like MSIK-1 and MSIK-2. The MSIK poses of the manipulator are illustrated in Figure 10.13. Note that the MSIK poses of the legs are the same as the PMCPLs (posture mode changing poses of the legs). The most typical and classical parallel manipulator configuration is shown in Figure 10.14. A manipulator of this kind is known with the generic name Stewart\u2013Gough platform. The configuration of such a manipulator is denoted as 6UPS, which means that the manipulator has six legs and each leg comprises one universal joint with the fixed platform, one prismatic joint between the lower and upper links of the leg, and one spherical joint with the moving platform. The manipulator is actuated through the prismatic joints. If hydraulic actuators are used, then the prismatic joints are replaced with cylindrical joints and the configuration of the manipulator is denoted as 6UCS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002184_icuas.2016.7502555-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002184_icuas.2016.7502555-Figure1-1.png", + "caption": "Fig. 1. Hexacopter - AscTec Firefly [16]", + "texts": [ + " The contents of the paper are as follows, the introduction is given here in Section I followed by the description of the multicopter system in Section II. The controller design which includes the inner-loop acceleration tracking controller and the adaptive augmentation is explained in Section III. Results and discussions are given in Section IV and the work is concluded in Section V. The necessary derivations and proofs for the stability are given in the Appendix. II. PLANT DESCRIPTION The multicopter used here is a hexacopter which is shown in Fig. 1 with its body-fixed axes defined by the B-frame. The translational states (position and velocity) are defined in the W -frame. The position kinematics are defined as( \u0307rR )W W = ( VR K )W W where ( rR ) W \u2208 R 3 and ( VR K )W W \u2208 R 3 are the position and velocity respectively. The translational dynamics are( \u0307VR K )WW W = ( aRK )WW W where the acceleration ( aRK )WW W \u2208 R 3 is given in terms of the specific propulsive force ( fRP ) B \u2208 R 3, specific aerodynamic force ( fRA ) B \u2208 R 3 and specific gravitational force ( fRG ) W \u2208 R 3 as ( aRK )WW W = MWB [( fRP ) B + ( fRA ) B ] + ( fRG ) W ", + " Therefore, the force and moment generated by the i-th propeller in the P - frame can be modeled as( ZPi P ) B,i =\u2212 kT\u03c9 2 i , (5)( NPi P ) B,i =\u2212 sgn (\u03c9i) kM\u03c92 i (6) which are the force and moment about its rotational axis respectively. Here kT \u2208 R and kM \u2208 R are rotor-specific force and moment coefficients and \u03c9i \u2208 R is the angular velocity of the i-th propeller. The plant parameters that are considered are given in Table I. Using (5)-(6) and from the spatial positions of each rotor illustrated in Fig. 1, the propulsive moments in B-frame are( MR P ) B = 6\u2211 i=1 [( rRPi ) B \u00d7 ( FPi P ) B,i + ( MPi P ) B,i ] , = 6\u2211 i=1 \u23a1 \u23a2\u23a3 \u239b \u239d l cos (\u03b1i) l sin (\u03b1i) 0 \u239e \u23a0\u00d7 \u239b \u239c\u239d 0 0( ZPi P ) B,i \u239e \u239f\u23a0 + \u239b \u239c\u239d 0 0( NPi P ) B,i \u239e \u239f\u23a0 \u23a4 \u23a5\u23a6 , =BM\u03c9u where ( rRPi ) B \u2208 R 3 is the position vector of the i-th propeller with respect to the reference point R, l \u2208 R is the length of the propeller arm and \u03b1i is the angle between xB and the rotor arm with the i-th propeller. The matrix BM\u03c9 \u2208 R 3\u00d76 is given by\u23a1 \u23a2\u23a3 0.5lkT lkT 0.5lkT \u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003172_14484846.2020.1769462-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003172_14484846.2020.1769462-Figure8-1.png", + "caption": "Figure 8. Boundary conditions for cam insertion with load for HD with proposed cam (A).", + "texts": [ + " There are no other contacts by tooth pairs beside the line of action. This is due to circular pitch contact around the line of action. Therefore, the full load of 194, N-m torque is applied at pitch points at 180\u00b0 apart along the major axis of deformed FG cup. Moreover, to get the combined effect of split-cam SWG insertion and applied load on FG cup, the FE simulation for insertion of split-cam SWG runs separately, and then the results of the simulation is fed to the FE simulation for applied load. The natural and essential boundary conditions are shown in Figure 8. The essential boundary conditions are same for each FE simulation, i.e. the displacement of closed end of FG cup is constrained in all direction motion. The stresses developed on FG cup due to combined effect in FE simulation is shown in Figure 9. The von-Misses stress developed on FG cup is shown. In experiments, the circumferential strain gauges are placed on three sections (at 30, 40 and 50 mm from open end of FG cup) on FG cup. However, longitudinal strain gauges are placed on two sections (at 30 and 50 mm from open end of FG cup) on FG cup" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.23-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.23-1.png", + "caption": "Figure 10.23 MSIK-1 of L1 viewed together with L3.", + "texts": [ + " However, since it belongs to an active joint, a suitable finite value can still be assigned to it, provided that the velocity of the end-effector (i.e. v) be specified so as to obey the following consistency condition that is implied by Eqs. (10.545) and (10.546). v\u22171c\ud835\udf19\u2032 k + v\u22172s\ud835\udf19\u2032 k = 0 (10.552) In other words, if the motion of the end-effector is planned consistently with Eq. (10.552), then it is not necessary to avoid MSIK-1(k) of Lk . This singularity, which is quite likely to occur with \ud835\udf03\u2032k = 0, is illustrated in Figure 10.23 by showing the first and third legs (again as if \ud835\udefd3 = \u2212\ud835\udf0b) when the first leg is in singularity. Incidentally, when Eqs. (10.445) and (10.441) are compared, it is seen that MSIK-1(k) also happens to be the PMCPL-k, i.e. the PMCPL of the leg Lk . On the other hand, Eqs. (10.548) and (10.550) imply that the leg Lk may get into a pose of MSIK-2(k) if c\ud835\udf19\u2032 k = 0, i.e. if \ud835\udf19\u2032 k = \u00b1\ud835\udf0b\u22152. In this singularity, which looks the same as the PSIK-2(k) of Lk , the lower links C\u2032 kA\u2032 k and C\u2032\u2032 k A\u2032\u2032 k get aligned with themselves and the chord C\u2032 kC\u2032\u2032 k of Lk " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002451_cacs.2015.7378369-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002451_cacs.2015.7378369-Figure5-1.png", + "caption": "Figure 5. Illustration of current reconstruction problem at the second sampled time: (a) Reference voltage vector in Sector I boundary region. (b) Corresponding three-phase PWM pattern. ( c) Voltage vectors for switching (proposed). (d) Corresponding three-phase PWM pattern (proposed).", + "texts": [], + "surrounding_texts": [ + "As illustrated in the previous section, Fig. Sea) and (b) presents a similar problem and its corresponding PWM signals, respectively. The problem occurs at the second sampled time tSGmp2 . This is because one of the output currents cannot be directly sampled if the duration of the active vector V; is insufficient. The distinguishing feature of the PWM signals waveform is two-short and one-long. Still using Sector I as an illustration, the second case is when the reference voltage vector Vre/ is close to an active vector V2 , which implies that half of the V; duration is possibly less than Tmm . This causes the current measurement validity at instant tsamp2 to be unreliable. Therefore, replace V; with V;' by extending the duration up to 2Trrun (by adding 1'1.\ufffd); a complementary voltage vector V4' during \ufffd = 1'1.\ufffd = 2T;run -\ufffd is added in the zero voltage to keep the average voltage vector unchanged. Fig. S( c) illustrates the detailed explanation. The same technique also ensures the PWM output signals are divided into two parts equally, and keep the waveform symmetrical in Fig. Sed). Furthermore, the add-on voltage vector V4' ensures the current measurement is valid during the two switching states and the sampling can be completed at the left side within one PWM interval. C. In the Case of Low Modulation Index (P3) At a low modulation index, the duty cycles of the PWM signals for the three phases will have an almost equal duration. The result is that the active voltage vector is not being used for long enough to ensure a proper sampling of the DC-link current. Fig. 6 illustrates the problem in a double-sided modulation strategy. Thus, two currents cannot be sampled correctly because V2 and V; are shorter than Tmm . It is necessary to generate a minimum time delay Tmin from which a proper sampling can be obtained. Taking both \ufffd and T2 as less than 2Trrun in sector I for example, both \ufffd and T2 are replaced by 2Tmin to ensure the phase current reconstruction detection demand is satisfied. Then, two complementary voltage vectors, V4' and Vs' are applied during \ufffd = 1'1.\ufffd = 2Tmm -\ufffd and T:, = I'1.I; = 2Tmm - I; , respectively. One approach to be addressed is to keep the minimum switching change such that the order should be \ufffd ---+ \ufffd' ---+\ufffd' ---+\ufffd' ---+ \ufffd. The method ensures the PWM output signals are divided into two parts equally and keep the waveform symmetrical. Furthermore, the add-on voltage vectors V4' and \ufffd' ensure the current measurement is valid during the two switching states and the sampling can be completed at the left side within one PWM interval. The proposed method adjusts the duty cycles within one switching period without changing the average voltage, to acquire the minimum required time by reducing the zero vectors. The amount of switching is increased by one or two times compared to the other adjusting scheme [8-11], which affects the THD and switching losses. However, because PWM signals are modified only if the active vectors are less than the minimum required time, the level of THD and increased switching loss do not seriously affect the system [10]. The proposed PWM adjusting technique and the related current reconstruction method can be unified as illustrated in Fig. 7. IV. SIMULATION RESULTS To verify the feasibility of the proposed method for a single-shunt current sensing technique, a current sensing scheme with an equivalent motor model was simulated via Matlab\u00ae Simulink software. The simulation results are derived from a permanent magnet synchronous motor (PMSM) with a DC-link voltage of 24 V, 14.4-kHz PWM carrier frequency, and 5 kHz current sampling frequency; Tm;n = 4.167 f.!s. The equivalent phase resistance and inductance are 1.623 Q and 2.625 mHo The back EMF coefficient\u00a2m of the PMSM is 0.0206 V(/rad/s). In general, if the THD of the phase current is less, then the motor torque ripple can be diminished. Because the back EMF of the PMSM is sinusoidal, to rotate the PMSM smoothly, choosing the motor phase current THD can be the performance criterion [18]. In this paper, the THD calculation is based on the Matlab\u00ae Power Graphical User Interface module (FFT) as one input to the base frequency. Furthermore, [19] discloses a common active vector insertion for Vo or V7 , and this approach is called the conventional adjusting scheme in this paper. From the simulated results shown in Table II, the proposed method of phase current THD is similar to the conventional adjusting scheme in the high-speed command. While in the medium-speed command, the proposed method is better than the conventional adjusting scheme. Furthermore, the proposed method is significantly superior to the conventional adjusting scheme in the low-speed command. The phase voltage THD of the proposed method is the worst one, which results in a small amount of switching loss due to the extra switching states required. TABLE n. Speed Command (Hz) 64 Hz (960 rpm) 40 Hz (600 rpm) 24 Hz (360 rpm) PHASE CURRENT/VOLTAGE THD IN DIFFERENT SPEED COMMANDS Phase-A Conventional Proposed THD Adiustin!! Scheme Method Current 2.09% 2.28% Voltage 87.71% 92.65% Current 5.88% 4.75% Voltage 140.19% 153.13% Current 28.00% 13.81% Voltage 207.62% 252.32%" + ] + }, + { + "image_filename": "designv11_34_0002251_b978-0-12-803581-8.03009-5-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002251_b978-0-12-803581-8.03009-5-Figure5-1.png", + "caption": "Figure 5 Schematic diagram of a remotely loaded hexagonal honeycomb and the geometric parameters associated with a typical cell.", + "texts": [ + " Analysis in the spirit of Gibson et al. (1982) is presented in this section, when plastic deformations are considered. Such abstractions to represent cellular materials are useful in that they provide valuable insight into the behavior of structured material in a simple and effective way. The model considered here is that of two-dimensional hexagonal honeycomb. A honeycomb lattice, loaded at infinity, undergoes plastic deformation when the stress within the cell wall exceeds the yield stress of the constituent material (Figure 5). Each point along the strut experiences a different state of stress. Some cross-sections may be fully elastic, while others may be partially plastic. Plastic zones spread along the struts, lengthwise as well as thickness wise, when the remote stress is increased. Due to mirror symmetry of the lattice about cell walls aligned to the direction of the remote stress, and due to the force equilibrium, the vertical members are stress free and only the inclined cell walls participate in the deformation. Therefore, the bulk material properties can be inferred from the response of the tilted struts. The cell walls can be adequately modeled as a cantilever beam when they are thin. The mechanical response of a lattice sheet is related to the strut response through the geometric parameters. For a hexagonal cell of edge of length l, height h and internal angle y (see, Figure 5), strain along and across the direction of the load application are given by: ejj \u00bc \u03b4tany l and e> \u03b4cosy \u00f02lsiny\u00fe h\u00de \u00bd10 where \u03b4 is the deflection of the tip of inclined members with respect to their roots, measured transverse to them. When the cell walls deform plastically, nonlinearities are introduced into the problem and the superposition principle does not hold which introduces mathematical difficulties in the analysis. To obviate this problem, the deformed shape of the struts is calculated by integrating the curvature accounting for the stress distribution along the cell walls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.13-1.png", + "caption": "Figure 9.13 Line diagrams that show the kinematic details.", + "texts": [ + " Indicating the order of the joints as well, the manipulator is designated symbolically as 2R-P-3R or R2PR3. The joint axes are also shown in the figures together with the relevant unit vectors. The prismatic joint 3 and the succeeding revolute joint 4 form a cylindrical joint arrangement. Moreover, the last three revolute joints (4, 5, 6) form a spherical joint arrangement. The kinematic details (the joint variables and the constant geometric parameters) of the manipulator are shown in the line diagrams in Figure 9.13. The line diagrams comprise the side view, the top view, and two auxiliary views that show the joint variables that are not seen in the side and top views. The significant points of the manipulator are named as follows: O: Center point or neck point (origin of the base frame) S: Shoulder point, R: Wrist point, P: Tip point (a) Rotation Angles \ud835\udf031, \ud835\udf032, \ud835\udf033 = \ud835\udeff3 = 0, \ud835\udf034, \ud835\udf035, \ud835\udf036 Five of the rotation angles are joint variables, which are shown in Figure 9.13. The third one is associated with the cylindrical arrangement of the joints 3 and 4. So, it is taken to be zero as discussed in Chapter 7. Thus, u\u20d7(3) 1 is arranged to be parallel to u\u20d7(2) 1 . (b) Twist Angles \ud835\udefd1 = 0, \ud835\udefd2 = \u2212\ud835\udf0b\u22152, \ud835\udefd3 = \ud835\udf0b\u22152, \ud835\udefd4 = 0, \ud835\udefd5 = \u2212\ud835\udf0b\u22152, \ud835\udefd6 = \ud835\udf0b\u22152 (c) Offsets d1 = 0, d2 = OS, s3 = SR, d4 = 0, d5 = 0, d6 = RP Five of the offsets are constant. The one associated with the prismatic joint is naturally variable. The variable offset and the nonzero constant offsets are given the following special names", + " r = d2C\u0302(0,2)u(2\u22152) 3 + s3C\u0302(0,3)u(3\u22153) 3 \u21d2 r = d2eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 e\u2212u\u03031\ud835\udf0b\u22152u3 + s3eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 u3 \u21d2 r = d2eu\u03033\ud835\udf031 u2 + s3eu\u03033\ud835\udf031 eu\u03032\ud835\udf032 u3 \u21d2 r = eu\u03033\ud835\udf031(d2u2 + s3eu\u03032\ud835\udf032 u3) \u21d2 r = eu\u03033\ud835\udf031[d2u2 + s3(u3c\ud835\udf032 + u1s\ud835\udf032)] \u21d2 r = eu\u03033\ud835\udf031[u1(s3s\ud835\udf032) + u2d2 + u3(s3c\ud835\udf032)] (9.127) Equation (9.127) can also be written as r = eu\u03033\ud835\udf031 r\u2032 (9.128) In Eq. (9.128), r\u2032 = u1(s3s\ud835\udf032) + u2d2 + u3(s3c\ud835\udf032) (9.129) Note that r\u2032 is the matrix representation of the vector r\u20d7 in the link frame 1(O), i.e. r\u2032 = r(1). In other words, Eq. (9.129) indicates the following coordinates of the wrist point R in the frame 1(O), which can also be written down directly by inspection referring to Figure 9.13. r\u20321 = s3s\ud835\udf032 r\u20322 = d2 r\u20323 = s3c\ud835\udf032 \u23ab \u23aa\u23ac\u23aa\u23ad (9.130) Kinematic Analyses of Typical Serial Manipulators 253 On the other hand, the expansion of the expression in Eq. (9.127) leads to the following coordinates of the wrist point R in the base frame 0(O). r(0) = r = u1r1 + u2r2 + u3r3 (9.131) r = (eu\u03033\ud835\udf031 u1)(s3s\ud835\udf032) + (eu\u03033\ud835\udf031 u2)d2 + (eu\u03033\ud835\udf031 u3)(s3c\ud835\udf032) \u21d2 r = (u1c\ud835\udf031 + u2s\ud835\udf031)(s3s\ud835\udf032) + (u2c\ud835\udf031 \u2212 u1s\ud835\udf031)d2 + (u3)(s3c\ud835\udf032) \u21d2 r = u1(s3s\ud835\udf032c\ud835\udf031 \u2212 d2s\ud835\udf031) + u2(s3s\ud835\udf032s\ud835\udf031 + d2c\ud835\udf031) + u3(s3c\ud835\udf032) (9.132) Equation (9.132) indicates that r1 = s3s\ud835\udf032c\ud835\udf031 \u2212 d2s\ud835\udf031 r2 = s3s\ud835\udf032s\ud835\udf031 + d2c\ud835\udf031 r3 = s3c\ud835\udf032 \u23ab \u23aa\u23ac\u23aa\u23ad (9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003363_j.mechmachtheory.2020.104027-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003363_j.mechmachtheory.2020.104027-Figure9-1.png", + "caption": "Fig. 9. Schematic drawing of geometric model [2] .", + "texts": [ + " Therefore, (c) was studied, which is close to the characteristic of (b), and the radius of curvature does not become 0 even when the bending angle \u03b1k is near 0. This characteristic is defined as the proposed curve. In summary, this research analysed, experimented, and compared the vibration and noise characteristics of two chains with the following pin cross sections: involute curve (a) and proposed curve (c). The vibration and noise characteristics of two chains with pin cross sections consisting of an involute curve (a) and proposed curve (c) were analyzed and compared using a geometric model created by Nakazawa et al. [2] . Fig. 9 shows a schematic diagram of the geometric model. This geometric model can analyze the chain chordal behavior by solving the static force balance of each part. In addition, the model was simplified by considering only the tight chord (i.e., the chord traveling from the output pulley to the input pulley). Each part was assumed to be rigid and the model does not consider deformation or frictional force. The origin of the model was the rotation center O of the input pulley. (a) to (e) in Fig. 9 show the units classified by position. (a) shows a unit in the input pulley, (b) shows a unit in the input pulley inlet, (c) shows units in the chord, (d) shows a unit in the output pulley outlet, and (e) shows a unit in the output pulley. The model is composed of m units in (c) and one unit at the other positions, for a total of m + 4 units, where m is an arbitrary integer. In the figure, the contact position and force of the unit involved in the calculation are described at each position. The number of the unit to be considered is represented by k , and k = 1,2,\u2026m + 4 in order from the unit in the input pulley" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000465_978-81-322-1665-0_50-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000465_978-81-322-1665-0_50-Figure1-1.png", + "caption": "Fig. 1 Mobile robot kinematic parameters", + "texts": [ + ", there is a pure rolling contact between the wheels and the ground, and also, there is no lateral slip between the wheel and the plane. The modeling in Cartesian coordinates is the most common use, and the discussion will be limited to modeling in Cartesian coordinates. The robot has two fixed standard wheels that are at both sides of mobile robot and one caster wheel that is attached to the front and is differentially driven by skid-steer motion. The two driving wheels are independently driven by two motors to acquire the motion and orientation. Both the wheels have same diameter \u20182r\u2019 (Fig. 1). The position of the robot in the 2-D plane at any instant is defined by the situation in Cartesian coordinates and the heading with respect to a global frame of reference. The kinematics model of this type of mobile robot is defined by the following equations [16]: _x \u00bc vcosh \u00f01\u00de _y \u00bc vsinh \u00f02\u00de _h \u00bc x \u00f03\u00de where x and y are the coordinates of the position of the mobile robot. h is the orientation of the mobile robot with respect the positive direction X-axis. v is the linear velocity, and x is the angular velocity (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000196_tmag.2011.2153410-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000196_tmag.2011.2153410-Figure1-1.png", + "caption": "Fig. 1. IPM motor.", + "texts": [ + " APPLICATION OF SIMPLIFIED SD-EEC METHOD FOR MOTORS In this section, the application of the simplified SD-EEC method to an IPM motor and to an induction motor is described and the usefulness of our techniques is verified. We also try The IPM motor is a synchronous motor and consequently the magnetic vector potentials vary with the half cycle periodicity in the stator. However, in the rotor, the magnetic vector potentials have the dc component caused by the permanent magnets. Therefore, the simplified SD-EEC method is applied to only the magnetic vector potentials in the stator every half cycle period of the power frequency and the method is not applied to those in the rotor because they do not satisfy (2). Fig. 1 shows the analyzed model of an IPMmotor [11] driven by the sinusoidal voltage source. The electric conductivity of the permanent magnet is 694 444 S/m. Table I shows the analysis conditions. The simplified SD-EECmethod is applied only in the stator at 180 , 360 , 540 , 720 , and 900 of the electrical angles, which are the half cycle period of the power frequency. Fig. 2 shows the calculated current waveforms, which are normalized by the RMS value of the current at the steady state. By applying the simplified SD-EECmethod at 180 of the electrical angle, the error of the calculated current is greatly eliminated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure6-1.png", + "caption": "Figure 6: Non-uniform deformation during free-state stiffness test rig-force measurements", + "texts": [ + " As it can be seen from the figure, although some difference is observed in BTF values of two tested seals, the results are comparable with each other. In an ideal case for an unpressurized seal, the slope of the BTF-\u2206R curve, which equals to free-state seal stiffness, should remain constant. However, as it can be seen from Figure 5, the seal stiffness continuously increases as radial interference increases, and the measured free-state BTF deviates from the representative line. The main reason for the nonlinear increase in BTF during free-state stiffness test rig measurements is the nonuniform deflection of bristles, which is visualized in Figure 6. As the metallic pad interference is introduced during the FSS-TR tests, the end bristles contact the neighboring bristles and deform them. As a result of the cascaded contact with side bristles, additional contributions to the measured force are observed due to bending and friction from neighbor bristles. This in turn causes the BTF results to increase as shown in the graphs. Since the contribution of side bristles increases with increasing bristle deflection, the slope of the BTF-\u0394R curve increases continuously" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003179_s12046-020-01330-4-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003179_s12046-020-01330-4-Figure15-1.png", + "caption": "Figure 15. (a) box pushing task, (b) pick and place task, and (c) image capture task.", + "texts": [ + " This is due to the paper\u2019s focus where task allocation can be studied not just the algorithm alone, but also including the communication framework for real live experiment testing. Therefore, the focus involved the overall framework which does not have precedent and hence no comparison is made. Experiments are performed to implement the developed framework in a heterogeneous multirobot system comprising of 4 to 5 Khepera IV robots with different sensing capabilities and additional grippers, on a 2D 12 by 6 grid world structure. Three types of tasks are used in the experiments shown in figure 15. Box pushing task involved pushing a box two steps away from box original location, whereas pick and place task involved retrieving an object and placing it at new location. Image capture task involved capturing image in the 4-cardinal directions of the target location. The model experiment is divided into two sets. Experiment 1 is made to solely show the implementation of task allocation algorithm in a controlled environment to show the framework operations and key features of the algorithm. Experiment 2 is made to show a more practical application of the developed framework, with other factors such as obstacles in the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure28-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure28-1.png", + "caption": "Fig. 28. Absolute position vectors of the toroidal model.", + "texts": [ + " Placing axes X \u03beR and X \u03beF on this point, one can reach points R and F by making a clockwise rotation \u03b7i around axis Z \u03be so that X \u03b7i points to R, F . The corresponding rotation matrix R \u03b7i is: i R \u03b7i = \u239b \u239d cos (\u03b7i ) sin (\u03b7i ) 0 \u2212 sin (\u03b7i ) cos (\u03b7i ) 0 0 0 1 \u239e \u23a0 , i = R, F . (58) These rotations and the local frames are shown in Figs. 26 and 27 . The absolute position vectors of the centers of the tubes, denoted by r P R and r P F , and the absolute position vectors of the contact points, r R and r F , are shown in Fig. 28 and can be computed from: r P R = r G 4 + r G 4 P R = r G 4 + R 41 R \u03beR \u239b \u239d b 0 0 \u239e \u23a0 , r P F = r G 7 + r G 7 P F = r G 7 + R 71 R \u03beF \u239b \u239d b 0 0 \u239e \u23a0 , r R = r P R + r P R P R = r P R + R 41 R \u03beR R \u03b7R \u239b \u239d a 0 0 \u239e \u23a0 , r F = r P F + r P F P F = r P F + R 71 R \u03beF R \u03b7F \u239b \u239d a 0 0 \u239e \u23a0 . (59) The next step is to obtain the holonomic and non-holonomic constraints. The addition of the two non-generalized coor- dinates \u03b7R and \u03b7F means that the multibody model is defined by 15 coordinates: q = ( x 2 y 2 z 2 \u03b12 \u03b22 \u03b32 \u03b852 \u03b832 \u03b865 \u03b843 \u03b876 \u03beR \u03beF \u03b7R \u03b7F )T ", + " (60) Since the number of coordinates has been enlarged, the number of constraints should also be increased so that the model remains with 5 degrees of freedom. On the one hand, as in the case of the multibody model with ring wheels, holonomic constraints arise from forcing the contact of the wheels with the ground, being the corresponding set: C h ( q ) = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d r R Z r F Z n \u00b7 t L R n \u00b7 t T R n \u00b7 t L F n \u00b7 t T F \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 0 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 , (61) where t L R , t L F are the longitudinal tangent vectors to the contact points, and t T R , t T F are the transversal tangent vectors, depicted in Fig. 28 . These are given by Eqs. (62) and (63) : t L R = R 41 \u2202 \u0304r 4 G 4 R \u2202 \u03beR , t T R = R 41 \u2202 \u0304r 4 G 4 R \u2202 \u03b7R , (62) t L F = R 71 \u2202 \u0304r 7 G 7 F \u2202 \u03beF , t T F = R 71 \u2202 \u0304r 7 G 7 F \u2202 \u03b7F , (63) where r\u0304 4 G 4 R and r\u0304 7 G 7 F are: r\u0304 4 G 4 R = R \u03beR \u239b \u239d \u239b \u239d b 0 0 \u239e \u23a0 + R \u03b7R \u239b \u239d a 0 0 \u239e \u23a0 \u239e \u23a0 = \u239b \u239d (b + a cos (\u03b7R )) cos (\u03beR ) \u2212a sin (\u03b7R ) \u2212(b + a cos (\u03b7R )) sin (\u03beR ) \u239e \u23a0 , (64) r\u0304 7 G 7 F = R \u03beF \u239b \u239d \u239b \u239d b 0 0 \u239e \u23a0 + R \u03b7F \u239b \u239d a 0 0 \u239e \u23a0 \u239e \u23a0 = \u239b \u239d (b + a cos (\u03b7F )) cos (\u03beF ) \u2212a sin (\u03b7F ) \u2212(b + a cos (\u03b7F )) sin (\u03beF ) \u239e \u23a0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002815_1045389x19898251-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002815_1045389x19898251-Figure1-1.png", + "caption": "Figure 1. Legged robot: (a) schematic, (b) photograph and (c) layout of SMA springs.", + "texts": [ + " A small legged deformable robot driven by PZTs (Smits and Ballato, 1994) and SMAs can realize linear, turning, and flexible motions suitable for completing various tasks, such as obstacle crossing, material transportation, object capturing, and so on. The simulation results show the feasibility of designing the required motion modes by setting the sequence of lifting and dropping different legs. Therefore, this robot with minimal structure composition has the characteristics of small size, light weight, low cost, multi-mode motion, and easy controlling. This robot is mainly composed of one cross-shaped polyvinyl chloride (PVC) body (0.25 mm in thickness), four SMA springs, and six piezoelectric bimorphs, as shown in Figure 1. Its mass and appearance size are 5.6 g and 70 3 60 3 30 mm, respectively, more parameters can be seen in Table 1. Piezoelectric bimorphs are selected as driving legs, whose vibration makes the robot walk in plane. SMA springs are utilized to lift and drop different legs, realizing the switching of multiple motion modes because they contract to deform the PVC body when heating. Figure 1(a) shows that six piezoelectric bimorphs (Sinocera Piezotronics, Inc., Yangzhou, China) are inserted and glued into the PVC body. L1\u2013L4 are parallel to each other at 75 angle with the ground, but L5 and L6 are vertical which are a little off the ground. Furthermore, four Ti49.2Ni50.8 SMA springs (Ti-Ni SMA factory, Guangdong, China) with one-way shape memory effect are connected to the fixed piles on both ends of PVC, made into the cross arrangement on the upper and lower surfaces of PVC. There are two sets of heating wires at the head and tail of longitudinal SMA springs (LS1 and LS2) to control L1\u2013L4, and one set of heating wires in transverse ones (TS1 and TS2) to control L5 and L6, respectively, as shown in Figure 1(c). The finite element method based on COMSOL Multiphysics software is utilized for vibration modal analysis of this robot. The boundary condition is that the roots of piezoelectric bimorphs glued to the PVC body are regarded as fixed constraints. The initial condition is that the driving voltage of 100 V0-p is applied to four inclined legs, excluding the vertical two. Figure 2(a) shows the calculated vibration mode based on piezoelectric coupling, in which the calculated resonant frequency of this first-order bending mode is 200 Hz, and the maximum total displacement of the leg tip is about 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000420_acc.2015.7171111-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000420_acc.2015.7171111-Figure1-1.png", + "caption": "Fig. 1. Terminal phase of flight illustrating angle of obliquity (AoO) and angle of attack (AoA)", + "texts": [ + " The platform concept involves a missile that is released from an aircraft compartment and proceeds to a target through two phases of flight. The first phase is approach, where the motor is off and body lift is utilized to maximize range. The next phase is engagement and begins with the ignition of a rocket motor. Towards the goal of interception, the system level objectives are to minimize target miss distance and engage head-on with the body orientation and velocity vector aligned with the oncoming target. These requirements are commanded by guidance laws requiring angle of attack (AoA) and angle of obliquity (AoO) regulation, Figure 1, at the time of collision. To achieve these objectives, the missile ultimately relies on a robust autopilot that enables tracking of the guidance commands with sufficient performance. Our control design approach is based on linearized mathematical models of the aircraft dynamics. Although, these models may not be accurate descriptions of the real-world system a stable and robust autopilot is nevertheless required 1 Research Mechanical Engineer, Weapon Dynamics and Controls Sciences Branch, Air Force Research Laboratory, Eglin AFB, Florida benjamin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003047_icit45562.2020.9067209-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003047_icit45562.2020.9067209-Figure3-1.png", + "caption": "Fig. 3. Stator slot harmonics flux lines, neglecting the rotor cage reaction. The proper electric load is imposed using air-gap current points.", + "texts": [ + " The specific losses distribution is not uniform in the teeth because of the It is worth to notice that an high losses density takes place not only in the rotor tooth tips, but also in the teeth body. Notice that the rotor is not synchronous and then, very slowly (the slip is low), the rotor teeth with high losses will be different, imaging that the rotor losses distribution rotates, in the rotor reference frame, at the speed \u03c9s \u00b7 s/p. In Fig. 2(b), the rotor iron losses harmonic components are shown; these losses are mainly due to the stator slot harmonics of order 24, 48, 72, etc. (k (Qs/p \u00b1 1)). In Fig. 3, an example of the stator slot harmonics flux lines is shown, neglecting the cage reaction. With the particular combination of stator and rotor slots considered, the harmonic fields have to penetrate deeply inside the rotor teeth to close their magnetic path; this consideration can explain why in Fig. 2(a), such an high losses density is placed in the rotor teeth body. It is easy to imagine that the rotor cage is able to reply to this flux, preventing to those flux lines to penetrate inside the rotor teeth", + " For this purpose, a first MS simulation with the rated on-load current is made, in order to store the saturation map. For the following simulations, in which only the slot harmonics are imposed, the permeability in each element is linearized around the on-load working point, for each element of the mesh. So an incremental disturbance will act along the tangent of the BH curve in the working point from the previous MS field solution. Let\u2019s consider, at first, the unskewed rotor. The harmonic fields, in Fig. 3, vary, inside the rotor teeth, at very high frequency, causing the rotor cage reaction, that tends to shield these fields. In Fig. 7(a), the slot harmonics flux lines, neglecting the rotor reaction are shown: they are able to penetrate deeply inside the rotor teeth causing iron losses. On the other hand, when the rotor cage reaction is considered, the flux map becomes the one in Fig. 7(b). The rotor currents, induced by the stator slot harmonics (see Fig. 7(c)) are able to almost completely shield those flux lines that tend to penetrate inside the rotor teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000395_s10409-015-0399-4-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000395_s10409-015-0399-4-Figure4-1.png", + "caption": "Fig. 4 a Hinged rod. b Rigid rod", + "texts": [ + "01 4.4 \u22120.5 The number in the table is the variations from that maximum velocity is 4.57 m/s, \u03b81 is 24.9\u25e6 and \u03b82 is 68.8\u25e6 velocity. Since the torso and long neck of the woodpecker are hinged individually, the globalmovement of system is similar to the whipping of a rope. Therefore, the mechanism of this so-called rope-whipping phenomenon will be discussed in more detail. To understand the rope-whipping phenomenon, the movements of a hinged rod are compared with those of a rigid rod, as shown in Fig. 4, in which P1 is the fixed end that is hinged on the ground, P2 is the free end, and P3 is the middle hinge of the hinged rod. \u03b21 and \u03b22 are the angles between the segments and the horizontal line. Suppose rotational springs G1 and G2 exist at hinges P1 and P3, respectively. \u03b2i is the angle of the rod relative to the horizontal line. At t = 0, \u03b21 = 0, \u03b22 = 0, and the preloads at G1 and G2 are \u03b20 and 0, respectively. Then both rods are released from their respective stationary positions, and the equations of the hinged rod are as follows: ( ( Jc + 5 4ml2 ) 1 2ml2 cos (\u03b21 \u2212 \u03b22) 1 2ml2 cos (\u03b21 \u2212 \u03b22) ( Jc + 1 4ml2 ) ) ( \u03b2\u03081 \u03b2\u03082 ) + ( + 1 2ml2 sin (\u03b21 \u2212 \u03b22) \u03b2\u03072 2 \u2212 G1 (\u03b20 \u2212 \u03b21) + G2 (\u03b21 \u2212 \u03b22) \u22121 2ml2 sin (\u03b21\u2212\u03b22) \u03b2\u03072 1 \u2212 G2 (\u03b21\u2212\u03b22) ) =0, (2) where m and l are the mass and length of rod P1P3 or P2P3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000882_dscc2013-3887-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000882_dscc2013-3887-Figure2-1.png", + "caption": "FIGURE 2. COORDINATE SYSTEMS AND FORCES/MOMENTS ON A QUAD-ROTOR FRAME", + "texts": [ + " A study comparing different kinds of PN techniques is presented next followed by results obtained via simulation studies and experiments. Quad-rotor is a type of UAV which is lifted and propelled using four rotors placed at the four diagonally opposite ends of the UAV. The rotor angular velocities are used as four inputs to the system to control the three translation and three rotational degrees of freedom. The motion and dynamics of the quad-rotor can be studied with the help of two coordinate systems: the world frame (W) and the body frame (B). The coordinate systems and free body diagram of quad-rotor are shown in Fig. 2. The world frame W is determined by axes Xw, Yw, Zw. The body frame is attached to the center of mass of the quad-rotor. Let us define forward direction as positive XB direction pointing from the center of mass to rotor 1. Rotor 2 is in the positive direction of the YB axis and rotor 3 is in the negative of the YB axis. ZB direction is perpendicular to the XB, YB plane. We use Z-X-Y Euler angle to model the rotation of the quad-rotor in the world frame. To get B coordinates from W, we first rotate by angle (yaw) along the Zw, then rotate by angle (roll) along the intermediate X-axis, and then finally rotate by angle (pitch) along the intermediate Y-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002900_j.engstruct.2020.110218-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002900_j.engstruct.2020.110218-Figure1-1.png", + "caption": "Fig. 1. Parameterised fold-line geometry.", + "texts": [ + " The lattermost technique is employed by Industrial Origami for low-force manual folding of sheet metal products [31,32] and is of core interest in this paper. It utilises a curved slit to reduce facet stress concentrations [33], enables a highly precise folded centre of rotation, and can be digitally fabricated using a parametric hinge definition comprising three geometric parameters, slit spacing as, radius ar and polar angle a [34], and one manufacturing parameter, kerf width ak, as shown in Fig. 1. Kinematic analysis of foldable structures typically assumes perfectly rigid bar or panel elements, connected by idealised zero-stiffness hinges [35,36]. For characterisation of realistic mechanical or structural behaviours, extended analysis techniques have been developed to consider non-rigid panels [37] and nonlinear hinge stiffness [38]. Such methods of analysis rely significantly on accurate calibration of panel [39,40] and fold-line stiffness [41,4]. Panel deformation behaviour is well understood in existing literature [42], however understanding of fold-line stiffness behaviours is very limited" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001394_mtsj.47.3.4-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001394_mtsj.47.3.4-Figure1-1.png", + "caption": "FIGURE 1", + "texts": [ + " Conversely, when the actuator increases the displaced volume, the vehicle becomes less dense than the surrounding water and ascends. The total vehicle mass, which remains fixed, is m \u00bc mrb \u00fe mp \u00fe mb Define a body-fixed, orthonormal reference frame centered at the geometric center of the vehicle (the center of buoyancy) and represented by the unit vectors b1, b2, and b3. The vector b1 is aligned with the longitudinal axis of the vehicle, b2 points out the right wing, and b3 completes the right-handed triad (see Figure 1). Define another orthonormal reference frame, denoted by the unit vectors i1, i2, and i3, which is fixed in inertial space such that i3 is aligned with the force due to gravity. The relative orientation of these two reference frames is given by the proper rotation matrix RIB, which maps free vectors from the body-fixed reference frame into the inertial reference frame. Let ei represent the standard basis vector for R3, where i \u00ce{1, 2, 3}, and let the character \u0302 denote the 3 \u00d7 3 skewsymmetric matrix satisfying a\u0302b \u00bc a\u00d7 b for vectors a and b. Then, in terms of conventional Euler angles (roll angle \u03c6, pitch angle \u03b8, and heading angle \u03c8), we have RIB \u03c6; \u03b8;\u03c8\u00f0 \u00de \u00bc ee\u03023\u03c8ee\u03022\u03b8ee\u03021\u03c6 whe r e e ( \u00b7 ) d eno t e s th e ma t r i x exponential. The origin of the body frame sits at the vehicle\u2019s center of buoyancy (CB), which remains fixed under our assumptions. The center of mass (CM) of the glider (neglecting the contribution of mp) is located at rrb (see Figure 1). Let the inertial vector X = [x, y, z]T represent the position vector from the origin of the inertial frame to the origin of the body frame. Let v = [u, v, w]T and \u03c9 = [p, q, r]T represent the translational and rotational velocity of the body with respect to the inertial frame, but expressed in the body frame. The kinematic equations are X \u00bc RIBv \u00f01\u00de R IB\u00bc RIB ^\u03c9 \u00f02\u00de In addition to the 6 degrees of freedom associated with the vehicle\u2019s translation and rotation, there are 2 degrees of freedom associated with the moving mass, which is modeled as a particle. To describe the pointmass position, we define a third orthonormal \u201cactuator\u201d triad {a1, ar, a\u03bc}, where a1 is parallel with b1 (see Figure 1). The vector ar points in the radial direction from the vehicle centerline through the point mass. The vector a\u03bc completes the right-handed frame. The proper rotation matrix RBP maps free vectors from this particlefixed frame to the body frame, RBP \u00bc 1 0 0 0 sin\u03bc cos\u03bc 0 cos\u03bc sin\u03bc 0 @ 1 A where \u03bc is the rotation angle of the moving mass about the longitudinal axis of the vehicle. Let \u223crp \u00bc rpxa1 \u00fe Rpar \u00bc rpx Rp 0 0 @ 1 A p denote the particle\u2019s position with respect to the body frame origin, expressed in the particle frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001489_1754337114536554-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001489_1754337114536554-Figure2-1.png", + "caption": "Figure 2. Scheme of the device for SCATToptoelectronic system testing: (a) general view of the installation from above; (b) SCATT target (1) while testing the sensor movement registration (2) horizontally and (c) SCATT target (1) while testing the sensor movement registration (2) vertically; 3 \u2013 directing railings.", + "texts": [ + " The methods of SCATT simulator testing were based on the calibration method, that is, on comparing actual sensor movements relative to the target with the registration of these movements with optoelectronic system. To ensure the accuracy of the movements, a special device was developed (Figure 1). SCATT sensor was mounted with special fittings to a sled device in a way that it has the ability to move. A main axis of the sensor is placed horizontally and perpendicular to the longitudinal directing railings of the sled (Figure 2). The height from the sensor to the floor equals 140 cm, and the distance between the sensor window and the plane of the sensor target is L = 914 cm, which corresponds to the SCATT terms of use and the competition rules.7,17 The sled position was determined with an accuracy of 0.05 cm. The SCATT target was mounted vertically and perpendicular to the main axis of the sensor at the standard height from the floor (140 cm).7 Thus, the centre of the target was at this same height from the floor as the sensor was. The sensor position depends on the position of the sled relatively to the directing railings of the device. For the metrological testing of vertical movements of the sensor, the target was rotated at an angle of 90 counterclockwise relatively to the axis, which was perpendicular to the target\u2019s plane and which was passed through the centre of the target (see Figure 2(c)). Twenty measurements were made for each of the axes of the target: horizontal and vertical (Table 1). The sensor was installed in such positions where the system registered centres and target rings from 10 to 1 on both sides from the centre. The position of the sled was recorded as well (coordinate of the centre of the target was 10 cm). The graphs in Figure 3 show a dependence of the actual distance of the sensor axis from the centre of the target (on a caliper scale) on the results of optoelectronic system registration in four directions: left, right, down and up" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002022_0020720915591582-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002022_0020720915591582-Figure3-1.png", + "caption": "Figure 3. The relation between the title angles of the plate and the angles of the actuators.", + "texts": [ + " Assume that there is no backlash or electric deformation in the gears, the work done by the load shaft equals to the work done by the motor shaft \u00bc 1 Kg m For each motor, the model is RmJm KmKg \u20acq1 \u00fe Kb \u00fe RmBm KmKg _q1 \u00bc U \u00f07\u00de where q1 can be the actuators angles qx or qy. The state of the model (4) q represents the position and the angles of the plate, while the control input are the torques of the motors. We need to establish the relation between the tilt angles of the plate ( x and y) and the actuators angles (qx and qx), see Figure 3. Because the bars \u2018\u2018VBx\u2019\u2019 and \u2018\u2018VBy\u2019\u2019 are always vertical, the movements of the two sides of the bars are the same Lx sin qx1\u00f0 \u00de sin qx2\u00f0 \u00de\u00bd \u00bc Px sin x1\u00f0 \u00de sin x2\u00f0 \u00de\u00bd Ly sin qy1 sin qy2 \u00bc Py sin y1 sin y2 at UNIVERSITE LAVAL on July 9, 2015ije.sagepub.comDownloaded from The angles of the plate are calculated as x \u00bc sin 1 Lx Px sin qx\u00f0 \u00de sin qx0\u00f0 \u00de\u00bd \u00fe sin x0\u00f0 \u00de n o y \u00bc sin 1 Ly Py sin qy sin qy0 \u00fe sin y0 n o \u00f08\u00de where x0 and y0 are initial angles of the plates, and qx0 and qy0 are the initial angles of the motors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000850_tie.2013.2289854-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000850_tie.2013.2289854-Figure5-1.png", + "caption": "Fig. 5. Vector diagram without iron losses.", + "texts": [], + "surrounding_texts": [ + "A relationship among motor current amplitude I , speed \u03c9, and displacement angle \u03b4mec has been identified to allow for different test conditions. An ideal motor model, with no iron and no mechanical losses, is here adopted, disregarding iron saturation too, to identify the aforementioned relationship together with the adopted reference frames." + ] + }, + { + "image_filename": "designv11_34_0001169_gt2014-26891-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001169_gt2014-26891-Figure4-1.png", + "caption": "Fig. 4 Coupled rotors with parallel misalignment", + "texts": [ + " (5) are modified as follows [ ] [ ][ ] [ ] [ ][ ] [ ] [ ][ ] [ ] [ ] ; ; ; T T e ee e e e T T e ee e e T M T T C T T k T a n d T Q (8) Assembling the modified element mass, damping, stiffness matrices and force vector given in eq. (8), the equation of motion of driven rotor is [ ]{ } [ ] { } [ ]{ } { }2 2 2 2 2 2 2M q C q K q Q+ + =&& & (9) 4 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use Due to parallel misalignment, an obliquity (\u03b3) of line joining centre of coupling halves is generated as shown in Fig. 4. This obliquity has not been considered in the present analysis. Similar to the equations of motion of Timoshenko beam element, the motion of coupling in XZ and YZ planes is assumed to be uncoupled. The stiffness matrix of the flexible coupling corresponding to either of the coupling half can be represented by following equation, 11 22 23 32 33 4544 54 55 000 0 000 000 0 0 0 0 0 0 z x x x y y y z Fk x Fk k Mk k y Fkk Mk k \u03b8 \u03b8 \u23a7 \u23ab \u23a7 \u23ab\u23a1 \u23a4 \u23aa \u23aa \u23aa \u23aa\u23a2 \u23a5 \u23aa \u23aa \u23aa \u23aa\u23a2 \u23a5 \u23aa \u23aa \u23aa \u23aa\u23a2 \u23a5 =\u23a8 \u23ac \u23a8 \u23ac \u23a2 \u23a5 \u23aa \u23aa \u23aa \u23aa \u23a2 \u23a5 \u23aa \u23aa \u23aa \u23aa \u23a2 \u23a5 \u23aa \u23aa \u23aa \u23aa\u23a3 \u23a6 \u23a9 \u23ad \u23a9 \u23ad (10) Here Mx is the moment in XZ plane and My is the moment in YZ plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002531_acs.langmuir.6b03237-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002531_acs.langmuir.6b03237-Figure9-1.png", + "caption": "Figure 9. (a) Charge injection after equilibrium has been reached using a large drop. (b) Electrowetting experiment using a small drop at a fast time scale. (c) Bulk charges and their countercharges in the electrode. (d) Remaining charges after (c) has been subtracted from (b).", + "texts": [ + " It decreases if the charge injection occurs predominantly from the liquid side, whereas it increases if the charges come mainly from the electrode side. In the derivation of the previous model, the charge density inside of the insulator was assumed to be laterally homogeneous. However, for a real experiment, charge injection (if happening) occurs only where the conducting liquid lies. Thus, the lateral homogeneity hypothesis corresponds to the following \u201cthought experiment\u201d: \u2022 A large drop of a conducting fluid is deposited and allowed to reach equilibrium under an applied voltage, as shown in Figure 9a. A given charge pattern is then created inside of the dielectric (injection only from the liquid side in this example) and previous derivations of charge densities at equilibrium apply. \u2022 The large drop is removed and a fresh small drop is deposited to perform an electrowetting experiment, as can be seen in Figure 9b. By applying the linear superposition principle, the situation depicted in Figure 9b can be considered equivalent to the superposition of those in Figure 9c,d. Figure 9c corresponds to the case of laterally infinite layers of injected charges in the dielectric bulk and their compensation on the electrode side. These charges will not induce any lateral forces onto the liquid drop and thus will not participate in the electrowetting effect. Applying the standard EWOD equation to Figure 9d leads to \u03b8 \u03b8 \u03c3 \u03c1 \u03b3 = + \u2212\u239c \u239f \u239b \u239d \u239e \u23a0 e V cos cos 2 2eq Y 0 liq eq ic (10) For a weak conductivity insulator, the standard EWOD equation is then modified as follows \u03b8 \u03b8 \u03b3 \u03c1\u0394 = \u0394 \u2212V V e Vcos ( ) cos ( ) 1 4eq 0 ic eq (11) It will be assumed that this modified EWOD equation will hold even in the case of a realistic experiment (single small drop). Although it might not quantitatively be the case, the effect should still be maintained: increased and decreased electrowetting when charge injection occurs from the electrode side and from the liquid side, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure10-1.png", + "caption": "Fig. 10 Cutting region when peripheral cutting edge does not take part in machining and yB<>: c \u00bc \u2212arccos yB r d \u00bc arccos yB r 8<: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f033\u00de (4) For IEi in part IV, where ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2\u2212 h\u2212Rc\u00f0 \u00de2 q > yB, and R1 <>: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f034\u00de (5) For IEi in part V, where yCR1, the cutting region can be divided into four different parts, as shown in Fig. 13. In this case, part I, part II, and part III are all the same as they are in Fig. 11. So the corresponding swept area si of the three parts can be calculated by Eqs. (31), (37), and (33). While part IV can be treated the same as part IV in Fig. 10, we can use Eq. (34). Based on the analysis above, swept area si of any IEi on the cutting edge can be calculated according to Eqs. (31) to (39). Although cutter wear cannot be avoided in turn-milling, it can be balanced on whole cutting edge if the cutting parameters are optimized, thereby prolong cutter life and reduce the cost on cutter. From Section 2, we know that swept area of any IEi on cutting edge can be calculated, and it can indicate the cutter wear. In addition, from Fig. 9, we can see the surface of the workpiece after machined is actually prismatic, and its side length only depends on angle \u03b1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000630_ilt-11-2011-0103-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000630_ilt-11-2011-0103-Figure1-1.png", + "caption": "Figure 1 Manufacturing scheme of the bush", + "texts": [ + " The first operation in pocket bearing machining consists of turning a cylindrical surface within the bearing block, which has been previously divided in two halves. The radius of such machined surface, the pocket, must be: Rc R cc (1) where R is the journal radius, and cc is the clearance of the cylindrical pocket surface. Then, to manufacture the elliptical bearing surfaces in the journal housing, a second turning operation is run, after the two halves of the bearing block have been split up and constrained at a distance d by means of suitable shims. Figure 1 shows the geometry of the two bearing block halves, kept at a distance POc d (shim thickness) before the second turning operation is performed. In the same figure, the contour of material removed by such operation is represented by dash-dotted lines. In all of the following figures, bearing clearance is amplified for the sake of clarity. Design and fast verification of pocket elliptical journal bearings Fabrizio Stefani Industrial Lubrication and Tribology Volume 66 \u00b7 Number 3 \u00b7 2014 \u00b7 393\u2013401 In Figure 1, if O1 and O2 denote, respectively, the centers of the bottom and top pads of the \u201celliptical\u201d part of the bearing and Oc is the center of the cylindrical pocket, the center of the turning operation (the location of spindle center) is coincident with both O1 and O2. The point O, which is set in the middle of the distance POc, is the origin of the reference system used to determine the offset s OO1 OO2 of the lathe spindle axis. The shim thickness is divided in two parts OcO1 and PO1 measured by the parameters e1 and e2, which are referred to as the eccentricities of the bottom and top pads, respectively", + " After the second machining, the two halves are joined together so that the top pad drops until its center O2 moves below Oc at a distance e2, while O1 remains in the same location, at distance e1 from Oc. The radius of the surface machined during the second operation is: Rp Rc t R cp (2) where t is the cut depth and cp the clearance of the elliptical bearing pads. The eccentricity e1 is an important parameter, as it rules the bottom pad profile that determines the load-carrying capacity of the bearing in nominal conditions (the bottom pad preload m1 e1/cp). Let denote the half-opening angle of the pocket. Figure 1 allows one to calculate e1 by means of simple geometrical considerations, as follows: e1 O1Oc O1H OcH Rp 2 Rc 2 sin2 Rc cos (3) or to express e1 as a function of manufacturing parameters by means of the following equation: e1 s d 2 (4) Similarly, the top pad eccentricity is given by: e2 d 2 s (5) As the same bottom pad eccentricity e1 can be obtained by means of several (infinite) couples of s and d according to equation (4), different bearing designs exist characterized by the same bottom lobe profile", + " For the purpose of the implementation of the numerical code, the bearing clearance distribution must be calculated. Figure 2, by showing the journal profile when its center overlaps Oc, illustrates the bearing clearance parameters. At the end of machining, the elliptical pad clearance cp, in agreement with equation (2), is equal to the horizontal clearance cH of the bearing, if higher-order infinitesimals (the lengths AB and A\u2019B\u2019) are neglected. Let cv1 and cv2 be the vertical clearances of bottom and top pad, respectively. By comparing Figure 1 to Figure 2 and by using equations (1) and (2), the following assembly constraint equations can be formulated: cv1 cc (6) cV2 cp e2 cc t e2 (7) For design A (e2 0), the resulting vertical clearance on the bottom and top pads are cV1A cc and cv2A cp cc t, respectively, i.e. clearance is higher in the top pad than in the bottom pad. In nominal working conditions, such additional clearance does not cause higher side loss, as the oil on the top pad, due to cavitation, flows only in the circumferential direction (Couette flow)", + " The deviations of the gap driving dimensions are specified in the bearing technical drawings by means of explicit (narrow) tolerances, rather than general tolerances, as they strongly influence lubrication performance and formation of the thin oil film in the micrometric gap. Due to manufacturing errors in the second machining operation, the actual cut depth at the top pad center tv2 tv2 (as it can be measured on the bearing) is greater than the theoretical value (the basic size) tv2 furnished by equation (9). Hence, its deviation tv2 is always positive. By neglecting the positioning error of the lathe spindle center with reference to the origin O (Figure 1), so that the offset s is exact, equations (4) and (5) yield: e1 e2 d 2 (11) where d, e1 and e2 are the deviations relevant to shim thickness and pad eccentricities. The deviation Rp of the pad radius depends on the positioning error of the lathe spindle center O2 due to the shim thickness deviation as well as the machining errors relevant to the first and second operations, which cause the deviations Rc and tv2, respectively: Rp Rc e2 tv2 Rc d 2 tv2 (12) From the gap driving dimensions, the remaining design parameters cc, cp, essential to performing the numerical analysis according to equation (10), can be easily found together with the relevant limit deviations by applying equations (1) and (2) in LMC and MMC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003676_ecce44975.2020.9235378-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003676_ecce44975.2020.9235378-Figure1-1.png", + "caption": "Fig. 1. 3D view of the actuator.", + "texts": [ + " Section IV presents the performance criteria related to the force density and the force ripple. From those criteria, the results of a bi-objective optimization are shown and then discussed. Section V presents the description of the test bench, the characterization of the winding prototype and its comparison with an existing actuator. 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 3615 Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 29,2021 at 10:13:24 UTC from IEEE Xplore. Restrictions apply. The actuator that is considered in this study is illustrated in Fig. 1. The stator is a double-sided topology with permanent magnets mounted on a U-channel yoke. Each north pole is facing a south pole on the other side. The moving part of the actuator is made of a three-phase distributed winding. In contrast with radial flux rotary machines with a PCB having an optimized shape [6], Nl independant PCBs separated by an insulation layer, are stacked one on top of each other. Each PCB is made of a double-sided copper clad with copper tracks printed on both side according to the layout presented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure2-1.png", + "caption": "Fig. 2. The 2-way coupled system model.", + "texts": [ + " c) Squiggly cooling duct covering potted end-regions. The cooling duct is integrated in to the flange of the motor. The three cooling duct designs (wave, spiral and end region squiggly) have been studied and compared by using the following models. A 3-D electromagnetic model with active motor components (stator and rotor core, coils and permanent magnets) is coupled with a 3-D thermal model of the electric motor with axis, casing, cooling duct with water, coil insulation, air within the casing and active components, as illustrated in Fig. 2. The thermal model thus includes liquid cooling and internal air flow (due to rotor rotation). The coupled models are used to calculate losses and temperatures which are transferred between models until convergence is achieved, allowing the analysis of performance and thermal behavior of 978-1-7281-9945-0/20/$31.00 \u00a92020 IEEE 867 Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 15,2021 at 16:48:44 UTC from IEEE Xplore. Restrictions apply. the motor for a set of operating points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003851_sensors47125.2020.9278934-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003851_sensors47125.2020.9278934-Figure3-1.png", + "caption": "Fig. 3. Simulated strain along the steel plate in COMSOL Multiphysics when the deformation element is loaded with 15 N. The strain shows the expected behavior of a load cell for measuring compressive forces. The strain gauges are applied at 4 mm and 36 mm resulting in an expected strain of roughly 105 \u00b5mm\u22121.", + "texts": [ + " 2) to achieve different connection points from base body to steel plate and in order to investigate the strain transmission depending on the exposure area. This way, three deformation elements are manufactured. These deformations elements are reworked by milling and M5 threaded holes are drilled for fastening to the measurement setup. A finite element analysis (COMSOL Multiphysics 5.5, Burlington, USA) is carried out to find a valid region for the strain gauges for a nominal load without exceeding the yield strength. The simulation showed the expected behavior with strain characteristic of such a load cell for compressive forces (Fig. 3). The high strain peaks result from high mechanical stress in the corner points. Therefore, we chose to apply the strain gauges at 4 mm and 36 mm at the top and bottom surface of the steel plate to remain in a valid region for our comparison to measurements. In this case a strain of 105 \u00b5mm\u22121 is expected. For this load case a maximum mechanical stress of about 130 MPa occurs in the corner regions, which is lower than the yield stress of the materials with 200 MPa. Considering only the steel plate results in a maximum mechanical stress of about 30 MPa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure9-1.png", + "caption": "Figure 9. Position and posture of the units.", + "texts": [ + " Pui and Pri are given by (5) and (6), respectively. (5) (6) Here, l is the length between the center of a unit and its tip, and lr is the length between the center of the roller B and the tip of a unit. When the robot is fixed, the vector normal to the unit N, which passes through Pui, crosses the center of the sphere, as shown Fig. 10 (a). Then, Pui satisfies (7). (7) When the robot is moving, both of the rollers of the magnetic adhesion mechanism cannot adsorb mechanistically onto spherical surfaces. Therefore, we assume that roller B (Fig. 9 (a)) constantly adsorbs onto surfaces. When the robot is moving, there is a certain distance between roller B and the sphere, as shown in Fig. 10 (b). Then, Pri satisfies (8). (8) Here, R is the radius of the sphere, and r is the radius of the roller. The motions of omnidirectional locomotion on spherical surfaces are generated from the above robot model. The locomotion strategy is schematized in Fig. 11. In this strategy, we define an imaginary wave propagation line. In previous research [12], we defined the propagation line as a straight line, because we expected the robot to move only on flat surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000465_978-81-322-1665-0_50-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000465_978-81-322-1665-0_50-Figure2-1.png", + "caption": "Fig. 2 Initial position of the robot", + "texts": [ + " The position of the robot in the 2-D plane at any instant is defined by the situation in Cartesian coordinates and the heading with respect to a global frame of reference. The kinematics model of this type of mobile robot is defined by the following equations [16]: _x \u00bc vcosh \u00f01\u00de _y \u00bc vsinh \u00f02\u00de _h \u00bc x \u00f03\u00de where x and y are the coordinates of the position of the mobile robot. h is the orientation of the mobile robot with respect the positive direction X-axis. v is the linear velocity, and x is the angular velocity (Fig. 2). ANFIS is one of the hybrid intelligent neuro-fuzzy system and it functioning under Takagi\u2013Sugeno-type FIS, which was developed by Jang [1]. ANFIS has a similar structure to a multilayer feed-forward neural network, but the links in an ANFIS only indicate the flow direction of signals between nodes, and no weights are associated with the links. There are two learning techniques are used in ANFIS to show the mapping between input and output data and to compute optimized of fuzzy membership functions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001094_gt2014-25395-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001094_gt2014-25395-Figure2-1.png", + "caption": "FIGURE 2. Finite Volume in a cylindrical coordinate system", + "texts": [ + " To determine the thermal parameters referred to table 1 it is necessary to know the temperature distribution inside the shaft in relation to the boundary conditions. The steady-state temperature distribution in a rotation-symmetric shaft can be calculated by solving the three-dimensional Laplace-equation in cylindrical coordinates (9) [9]. 0 = \u2206T = \u2202 2T \u2202 r2 + 1 r \u2202T \u2202 r + 1 r2 \u2202 2T \u2202\u03d52 + \u2202 2T \u2202 z2 (9) To solve this partial differential equation the Finite-VolumeMethod is used. Therefore the shaft is divided into a finite number of volumes according to Fig. 2 with as constant assumed temperature each. To reduce the order of the differential equation by one the integral form of Eq. (9) is transformed using the divergence theorem. \u222b \u2206T dV = \u222b \u2207T \u00b7ndS (10) Replacing the integral of the temperature-gradient over the surface by a sum of differences leads to one linear equation for every single volume: \u222b \u2207T \u00b7ndS \u2248 cu \u00b7Tu + cd \u00b7Td + cl \u00b7Tl+ cr \u00b7Tr + c f \u00b7Tf + cb \u00b7Tb\u2212 cC \u00b7TC , (11) with cu = \u2206\u03d5\u00b7ra\u00b7\u2206z \u2206ru , cd = \u2206\u03d5\u00b7ri\u00b7\u2206z \u2206rd , cl = \u2206r\u00b7\u2206z rmi\u00b7\u2206\u03d5l , cr = \u2206r\u00b7\u2206z rmi\u00b7\u2206\u03d5r c f = \u2206\u03d5\u00b7rmi\u00b7\u2206r \u2206z f , cb = \u2206\u03d5\u00b7rmi\u00b7\u2206r \u2206zb , cC = cu + cd + cl + cr + c f + cb The implementation into a linear system of equations results in: K Cond \u00b7T = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001090_icems.2014.7014015-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001090_icems.2014.7014015-Figure1-1.png", + "caption": "Fig. 1. Configuration of the DSBDFIG-WPGS (a) System configuration. (b) Schematic of the DSBDFIG", + "texts": [ + " Then in Section III, the model of wind turbine and mathematical model of the DSBDFIG are given. Decoupled vector control of the DSBDFIG is presented with detailed mathematical derivation in section IV, where a simple tip speed ratio (TSR) maximum power point tracking (MPPT) control method is also adopted. Finally, simulation results are presented in Sections V to validate stable operation of the DSBDFIG-WPGS and good dynamic performance of the proposed control scheme. II. BRUSHLESS DOUBLY-FED WIND POWER GENERATION SYSTEM A. System topology Fig. 1(a) shows the configuration of the DSBDFIG-WPGS. The wind turbine extracts wind energy and converts it into mechanical energy. The drive train (including shaft, bearing, and gearbox) subsequently transfers mechanical energy to the DSBDFIG, which converts mechanical energy into electric energy with the help of a back-to-back bidirectional converter. This DSBDFIG-WPGS is similar to the conventional DFIG-WPGS except for the generator itself. As shown in Fig. 1(b), the DSBDFIG consists of two stators and a cup-shaped rotor with closed windings. The balanced three-phase outer stator winding set is connected to the grid, while the balanced three-phase inner stator winding set is supplied with a grid connected bidirectional converter which only handles a fraction of the rated power. The specialized rotor has duallayer cores and each has a balanced three-phase winding set, and the two rotor winding sets are connected in reverse sequence to produce additive electromagnetic torque [21]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002546_978-3-319-47500-4_4-Figure4.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002546_978-3-319-47500-4_4-Figure4.6-1.png", + "caption": "Fig. 4.6 aA schematic of the follicle sinus complex (FSC). bA simple lumped parameter model of the whisker-FSC system. The viscoelastic behavior of the FSC neck is modeled by a torsional spring and damper arranged in parallel with stiffness k1 and damping coefficient b1. The boundary conditions at the HB are modeled by a linear spring and damper arranged in parallel with stiffness k2 and damping coefficient b2. Together, these end conditions provide a simplified representation of the FSC mechanical structure [16]. Reprinted with permission from IEEE. Copyright (2013) IEEE", + "texts": [ + " The coin head (CH) was designed at the base of the whisker for optimal transduction and sensing of whisker shaft movements (Figs. 4.4 and 4.5). The follicle is the part of a pinniped muzzle that holds the whisker shaft. The follicle structure, known as follicle sinus complex (FSC), is where all the mechanoreceptors that convey whisker signals into the nervous system reside. Design of the follicle is equally important to that of whisker shaft in achieving the high signal-to-noise ratio (SNR). Valdivia y Alvarado et al. developed a simple FSC model for the whisker-inspired sensor [15, 16]. Figure 4.6a shows a simplified diagram of the major components of follicle sinus complex. The whisker shaft (WS) is embedded inside the follicle and zones of innervation span the entire length (l) of the follicle. The neck of the follicle is composed of the outer conical body (OCB) and the inner conical body (ICB) which contain fine caliber innervation and a sebaceous gland. The upper half has an open lumen known as the ring sinus (RS). The lower half is referred to as the cavernous sinus (CS). The glassy membrane (GM) and the mesenchymal sheath (MS) surround the WS. The WS ends at the hair papilla (HP) which is surrounded by the hair bulb (HB). These anatomical structures not only contain the nerve endings required for transduction of mechanical signals of WS but they also provide a mechanical structure that supports the WS and influences its dynamic response to external perturbations. As a result, the follicle material properties must be designed carefully for optimal flow measurements. Figure 4.6b displays a simplified lumped parameter model of the whisker-FSC system designed by Valdivia y Alvarado et al. [16]. The FSC tissue has an average modulus EFSC of 10 kPa [19]. Figure 4.7a shows the basic whisker sensor design. A whisker of length L + l and inertia J is supported by a sensor module of length l which emulates the FSC. The sensor module has a thick but flexible viscoelastic membrane of modulus E1 and viscosity l1 that supports the whisker shaft and constrains its motions. Three different materials were chosen for the viscoelastic membrane to span a range of known FSC tissue properties and test the model\u2019s validity within that range", + " The flexible sensors consist of a coated substrate that displays changes in electrical conductivity as it is bent [20]. Each flexible sensor is mounted on a voltage divider circuit with a fixed DC voltage input (5 V). Deflections cause changes in resistance, which can be measured in the voltage drops across each sensor. According to the model developed by Valdivia y Alvarado et al. [16], the bending deflections hw of the whisker tip relative to its base can be modeled by a spring-inertia-damper system (with spring constant k, damping coefficient b, and inertia J) shown at the right of Fig. 4.6a. The FSC support of the whisker shaft can be modeled by a set of springs (with spring constants k1 and k2) and dampers (with damping coefficients b1 and b2) as shown in Fig. 4.6b which allow rigid body deflections h at the whisker base [16]. The whisker stiffness k is orders of magnitude larger than the FSC\u2019s combined stiffness and the whisker shaft can be assumed to oscillate as a rigid body (hw = 0). When excited by distributed forces F (z, t) due to fluid interactions in a given plane the whisker-FSC dynamics can be modeled by a transfer function of the form, h s\u00f0 \u00de \u00bc L=2 Js2 \u00fe b1 \u00fe b2\u20182\u00f0 \u00des\u00fe k1 \u00fe k2\u20182\u00f0 \u00deF s\u00f0 \u00de; \u00f04:1\u00de where the inertia term J includes added mass effects" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000678_robio.2014.7090419-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000678_robio.2014.7090419-Figure2-1.png", + "caption": "Fig. 2: Illustration of ades and a for relative field of view.", + "texts": [ + " \u03c3 13(\u03b7 b,q1) = yee1 \u2212 yb (14) \u03c3 13,d =C (15) Similarly to Task 1, the Jacobian is derived by taking the time derivative of \u03c313 using (2) and (1). \u03c3\u0307 13 = J13(\u03b7 b,q1)\u03b6 1 (16) Manipulator 2 - Task 1 - Relative Field of View - 1 DOF: End effector 2 should always point towards end effector 1. In this case, the task is chosen as the error between the desired and actual direction of the manipulator, so the desired task value is 0: \u03c3 21(\u03b7 b,q1,q2) = \u221a (ades \u2212a)T (ades \u2212a) (17) \u03c3 21,d = 0 (18) where ades and a are unit vectors illustrated in Figure 2 and are defined as ades = 1\u221a (xee1 \u2212 xee2)2 +(yee1 \u2212 yee2)2 \u23a1 \u23a3xee1 \u2212 xee2 yee1 \u2212 yee2 0 \u23a4 \u23a6 \ufe38 \ufe37\ufe37 \ufe38 pe (19) a = \u23a1 \u23a3cos(\u03c8ee2) \u2212sin(\u03c8ee2) 0 sin(\u03c8ee2) cos(\u03c8ee2) 0 0 0 1 \u23a4 \u23a6 \u23a1 \u23a31 0 0 \u23a4 \u23a6 . (20) The corresponding 1\u00d72 Jacobian is defined in [3] as \u03c3\u0307 21 = J21(\u03b7 b,q1,q2)\u03b6 2 (21) J21 = (ades \u2212a)T \u2016ades \u2212a\u2016 (\u2212S(ades)S(pe) \u2020J p +S(a)Jo ) (22) where S(\u00b7) is the matrix cross product operator and X \u2020 denotes the Moore-Penrose inverse of the matrix X . J p and Jo denote the position and orientation Jacobian matrices of end effector 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002258_itec.2016.7520251-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002258_itec.2016.7520251-Figure3-1.png", + "caption": "Fig 3. Scheme of an equivalent IPMSM with internal short circuits.", + "texts": [ + " IPMSM EQUATIONS UNDER INTERNAL SHORT CIRCUITS Voltage equations are shown for both the total voltage and the shorted turns respectively. = + (17) = , + (18) where, = 0 00 00 0 ; = 0 00 00 0 (19) and rs is the resistance for each winding without faults. To find the torque, (9) is used to calculate the co-energy in the windings, then the torque expression is given by: = 2 12 ( ) + ( ) (20) Based on (9), (12), (17) and (20) the model of the machine can be represented by the equivalent current and two equivalent currents , , , flowing through the healthy and shored turns respectively (Fig. 3). The flux linkage and torque behavior depends mainly on . As shown in (14) the equivalent current have embedded the failure path current, thus (13) and (18) must be used to obtain , . The only currents available to measure are the ones in the healthy turns, thus it is convinient to represent the model of the IPMSM as shown in Fig. 4, which is transformed to the rotor reference frame The matrix \u2126 include the rotor speed in electrical rad/s and it\u2019s expression is given by: \u2126= 0 00 \u2212 00 0 0 (21) The rest of the matrices in the model are obtained using the transformation matrices ( ) and ( ) given by [9]: ( ) = cos( ) cos \u2212 cos \u2212sin( ) sin \u2212 sin \u2212 (22) ( ) = cos( ) sin( ) 1cos \u2212 sin \u2212 1cos \u2212 sin \u2212 1 (23) thus: ( ) = ( ) ( ) (24) ( ) = ( ) + ( ) (25) = ( ) ( ) ( ) (26) , ( ) = ( ) , ( ) ( ) (27) , ,, ( ) = ( ) ,, ( ) ( ) (28) ( ) = ( ) ( ) (29) And finally, the electrical torque equation is given by: = 32 2 \u2212 \u2212 \u2212 (30) Based on this derivation, a complete model of the machine can be developed to have a strong understanding of the motor under the short circuit condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure11-1.png", + "caption": "Fig. 11. (a) The maximum shear stress on the shaft; (b) Factor of Safety of shaft.", + "texts": [ + " The material taken was IS2062 Grade E250. The stresses developed is shown in Fig. 10 (a). From Fig. 10 (b), it can be seen that the connecting rod is safe to operate under the given conditions. Also, the results are similar to what was obtained through hand calculations in Section 2.2. The analysis of shaft was done by taking supports and forces as shown in Fig. 2 and Fig. 3. The material taken was AISI 4130. The results obtained are for gradual loading. The shear stresses developed is shown in Fig. 11 (a) and the factor of safety in Fig. 11 (b). It can be seen that the shaft will not fail under the given conditions and also, the results are similar to what was obtained through hand calculations in Section 2.4. Please cite this article as: A. Sinha, S. Mittal, A. Jakhmola et al., Green energy generation from road traffic using speed breakers, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.06.248 In this paper a modern day speed breaker has been designed that proposes a method of power generation by harnessing the power of moving vehicles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000169_978-94-007-6046-2_14-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000169_978-94-007-6046-2_14-Figure4-1.png", + "caption": "Fig. 4 Two DOF cylindrospheric module used in the neck, shoulder, and elbow of NAO", + "texts": [ + " In addition, this structure helps better distribute the power between the hip roll joint and the pelvis joint and creates a specific motion style to the NAO humanoid. If the very first prototypes of NAO were based on servo motors; the integration of these actuators would not be compatible with the very strong constraints that Aldebaran aimed for on the physical appearance of the robot. For this reason, specific actuators, based on DC motors, gears, and position sensors were created. 10 degrees of freedom in NAO rely on the same type of actuation module called the cylindrospheric module (Fig. 4). This module includes two motors that provide two perpendicular motions. This module is very compact and makes the integration of motorization very easy. This module is used in the shoulder, at the elbow, and in the neck of NAO. This modularity is an important aspect to consider when regarding the goal of mass production for NAO. Another constraint that led to the design of proprietary actuators was research in backdrivability. Classically, humanoid robots use harmonic drives which provide high reduction ratio in a small volume, but are not backdrivable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000736_s00419-014-0845-y-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000736_s00419-014-0845-y-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of flywheel rotor stress analysis", + "texts": [ + " In this paper, the stress analysis of the anisotropic flywheel rotor under a high-speed rotation is conducted based on plane stress assumption, and the maximum radial stress location equation is established by extreme point method, which is solved by Newton iteration method. Specially, the effect of the anisotropy degree of the flywheel rotor material, the ratio of inner and outer radii and the rotation speed on the maximum radial stress location are systematically studied. The composite flywheel rotor with circumferential winding structure can be regarded as the orthotropic plane problem [18]. Therefore, it is convenient to establish a basic modeling for analysis in the cylindrical coordinate system, as shown in Fig. 1, r represents the axis along the radial direction and \u03b8 denotes the axis along the circumferential direction (filament winding direction); z is the axis direction coincides with the flywheel. The flywheel rotor in the paper is in a disc shape with a uniform distribution of stress along the thickness direction, which can be the axisymmetric plane stress state. In the analysis, the composite flywheel rotor is assumed to remain at a constant increased temperature, and all boundary conditions are independent of time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001340_amr.1004-1005.1344-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001340_amr.1004-1005.1344-Figure1-1.png", + "caption": "Fig. 1 Contact region of axial ring rolling 1-ring 2-conical roller", + "texts": [ + " Rolling force is always an important parameter for researching the ring rolling technology, which is related to determining the technical parameters of machine and designing structure to use reasonably. It is also involved in the selection of motor and the design of hydraulic system. By understanding the relationship and change rule between rolling force and the process parameters of ring rolling, The metal flow can be analyzed roundly and the amount of deformation can be controlled [4]. The contact region of axial ring rolling is shown in Fig. 1. The profile curve of the contact surface is the intersecting line where roll surface and workpiece end face. In radial-axial ring rolling process, the contact surface is a part of the surface of the conical roller. S is the feed of the conical roller. In order to facilitate the calculation of the contact area, the contact surface can be approximated as plane because S is much smaller than R that is the outer diameter of the workpiece. Establish the rectangular coordinate system that the origin is the center of ring, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001730_978-3-319-18296-4_13-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001730_978-3-319-18296-4_13-Figure4-1.png", + "caption": "Fig. 4 Wheel flat", + "texts": [ + " (29) and (35), the reaction of the rail results as N \u00bc Q\u00fe mV2 d 2yw dx2w \u00f037\u00de or N \u00bc Q\u00fe mV2 dxw dh d2 yw dh2 d2 xw dh2 dyw dh dxw dh 3 ; \u00f038\u00de if considering Eq. (30). It can be noticed that the rail reaction has two components \u2014one equals the wheel load and the other is proportional to the wheel mass and increases with V2. Finally, by substituting N from Eq. (36), we have the equation J d2h dt2 \u00bc Q\u00fe mV2 dxw dh d2 yw dh2 d2 xw dh2 dyw dh dxw dh 3 2 64 3 75q sin b; \u00f039\u00de which needs a solution. Here, the previously presented theory applies in the case of a wheel flat of radius R, which has the length of the flat defect 2l and the depth e (Fig. 4). The mathematical shape of the flat defect can be given as Fig. 3 The model of a wheel with irregular contour rolling along a smooth rigid rail where \u03b8o = arcsin(l/R). Next, the following values of the wheel parameters are to be taken into account: V = 10 m/s, R = 460 mm, l = 30 mm, e = 0.35 mm, m = 750 kg, J = 120 kg m2 and the wheel load Q = 100 kN. Figure 5 shows the flat defect contour and the wheel ideal contour. When using Eqs. (13) and (14), the trajectory of the wheel center can be computed, (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003932_s40962-020-00549-5-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003932_s40962-020-00549-5-Figure5-1.png", + "caption": "Figure 5. Temperature field result of fluid: a No. 1, b No. 10, c No. 20, d No. 27.", + "texts": [ + " In the present work, as shown in Figure 4, five main dimensional parameters for determining streamline type at the output are selected and simulation analysis is carried out changing these parameters as shown in Table 1. If we use the traditional experimental method, we have to make 35 = 243 experiments. Thus, in this work we made 27 simulation experiments using L27 (313) orthogonal table as shown in Table 219. At this time simulation experiments were made in SolidWorks Flow Simulation. As shown in Table 2, S/N ratio was calculated by the following equation because the lower the temperature is at the output, the better the characteristic is. g \u00bc 10 lg y2 Eqn: 1 Figure 5 shows the result of simulation experiments No. 1, 10, 20 and 27. As shown in Table 2, in No.17 experiment the lowest temperature at the output is - 189.26 C, which is the best. Based on the result of simulation experiments, we do optimization using simulation annealing. First, a prediction model is made using secondary multiple regression analysis. The model which predicts minimum temperature at the output is as follows: y \u00bc 1:4965x1 2:4922x2 \u00fe 0:027447x2 1 0:0004113x1x2 0:0016444x1x3 0:00048333x1x4 \u00fe 0:059556x1x5 \u00fe 0:0037906x2 2 0:019381x2x3 \u00fe 0:01541x2x4 \u00fe 0:25794x2x5 \u00fe 0:052841x2 3 \u00fe 0:0097348x3x4 0:0029764x2 4 0:59321x4x5 Eqn: 2 where y denotes the minimum temperature at the output, C; x1 up-profile radius 1, mm; x2 up-profile radius 2, mm; x3 down-profile radius 1, mm; x4 down-profile radius 2, mm; x5 thickness, mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002569_iceee.2016.7751210-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002569_iceee.2016.7751210-Figure1-1.png", + "caption": "Fig. 1. Description of roll control system [6].", + "texts": [ + " The yaw motion is achieved by deflecting a flap on the rudder, and sideslip motion is entered by lowering a wing and applying exactly enough opposite rudder so the aircraft does not turn. Yb and Zb represent aerodynamics force components, \u03d5 and \u03b4a represent the orientation of the aircraft in the earth-axis 978-1-5090-3511-3/16/$31.00 2016 IEEE system and aileron deflection angle respectively. Fig. 2 shows the forces, moments and velocity components in the body fixed frame of an aircraft. Roll control system is shown in Fig. 1, where L,M and N represent the aerodynamic moment components, the term p, q and r represent the angular rates components of roll, pitch and yaw axis and the term u, v and w represent the velocity components of roll, pitch, and yaw axis. In the model devlopement, we asume that the aircraft is in steady cruise with constant altitude as well as constant velocity, that a change in the pitch angle does not change de the speed of the aircraft and that the reference flight conditions are symmetric with propulsive forces constant [8] [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003402_s11071-020-05852-8-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003402_s11071-020-05852-8-Figure9-1.png", + "caption": "Fig. 9 Physical realization of the nonlinear device", + "texts": [ + " The nonlinear spring is designed in such a way that it can be mounted on any flat surface of appropriate dimensions. A mounting bracket is used to attach the device to any base structure with a flat face, on any configuration. This bracket also works as a mounting vise to fix the device to the rigid surfaceswhile applying pressure to the beam to maintain a tight cantilevered boundary condition. The device is fabricated in a CNC machine at the Mechanical EngineeringMachine Shop and is shown in Fig. 9. The spring assembly is mounted on a six-DOF servo hydraulic shake table manufactured by Shore Western Manufacturing, controlled by a SC-6000 PID-type servo controller at each DOF. The shake table dimensions are 760 \u00d7 760mm, with a maximum payload capacity of 200kg. The acceleration transducers are tightly glued to the mass on the tip of the beam, for recording the acceleration of the motion in both directions of the trajectory. Because the actual motion of the mass travels on a curved trajectory, and the transducers used are single direction, it is not possible to directly record measurements in the rectilinear directions (u and v), rather accelerations are recorded in normal and tangential coordinates, which are later corrected with the appropriate angle to horizontal and vertical components (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002642_phm.2016.7819811-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002642_phm.2016.7819811-Figure3-1.png", + "caption": "Figure 3. Tooth modeling: (a) a perfect gear tooth model reported in [16] and (b) tooth model with a single pit", + "texts": [ + " A paper documenting detailed results has been submitted to a journal for possible publication [18]. II. MESH STIFFNESS DERIVATION FOR GEARS WITH TOOTH PITTING In [11, 12, 15], the gear system is assumed to be without friction, manufacturing error, and transmission error; and the gear body is treated as rigid. The same assumptions are employed in this paper. Our goal is to express the mesh stiffness equations as a function of gear rotation angle. Toth In [16], a cantilever beam model was used for a gear tooth to evaluate the mesh stiffness of perfect gears as shown in Fig. 3 (a). The gear tooth was modeled starting from the root circle. The tooth fillet curve is not an involute curve and hard to describe analytically. Therefore, straight lines UU' and VV' are used to simplify the tooth fillet curve as did in [16] for the convenience of equation derivation. These two lines are given on Fig. 3. We extend this model by adding tooth pitting. In this section, a single circular pit is considered as shown in Fig. 3 (b). The whole circle is within the tooth surface area. The position and size of a single pit can be fully expressed by three variables: ( u , r, \u03b4 ), where u represents the distance between the tooth root and the circle center of the pit, r is the radius of the pit circle, and \u03b4 is the pitting depth. According to the involute curve properties, the action line of two meshing gears is tangent to the gear base circles and normal to the tooth involute profile. The action force F which is along the action line, can be decomposed into two orthogonal forces aF and bF , as shown in Fig", + " The expression of 2\u03b1 and 3\u03b1 are given as follows [16]: 2 0 0tan 2Z \u03c0\u03b1 \u03b1 \u03b1= + \u2212 , (12) 2 3 sinarcsin( )b r R R \u03b1\u03b1 = , (13) where Z is the number of teeth of the external gear, and 0\u03b1 is the pressure angle. For a gear tooth with pitting, the expressions of h , xh , xI and xA are different from the ones given above for a perfect gear tooth. In addition, the tooth contact width is not constant L. We use xL\u0394 , xA\u0394 and xI\u0394 to represent the reduction of tooth contact width, area and area moment of inertia of the tooth section, respectively, where the distance to the gear contact point is x. For a gear tooth with a single pit as modeled in Fig. 3 (b), the expressions of xL\u0394 , xA\u0394 and xI\u0394 are given as follows: 2 22 ( ) [ - , ] 0 others x r u x x u r u rL \u2212 \u2212 \u2208 +\u0394 = \uff0c \uff0c , (14) [ - , ] 0 others x x L x u r u r A \u03b4\u0394 \u2208 + \u0394 = \uff0c \uff0c , (15) 2 3 ( / 2)1 + [ - , ] 12 0 others x x x x x x x A A hL x u r u r I A A \u03b4\u03b4 \u0394 \u2212 \u0394 \u2208 +\u0394 = \u2212 \u0394 \uff0c \uff0c . (16) Given a gear tooth with a circular pit, we can calculate xL\u0394 , xA\u0394 and xI\u0394 for any tooth section of which the distance to the gear contact point is x. In addition, we can observe from (9) that x is a function of gear rotation angle (denoted by \u03b1 )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002814_lra.2020.2965914-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002814_lra.2020.2965914-Figure8-1.png", + "caption": "Fig. 8. Left: Section view of the particle jamming backed suction palm schematic. Right: Real prototype mounted on hand. The fully sealed system shown prevents uptake of food particulates and unwanted moisture with its octopus inspired, non-aspirating surface structures.", + "texts": [ + " It is suspected that pillars with rounded, as opposed to flat, tops may be more successful at expelling fluid into the channels because they contact the surface as a point rather than a plane. When the pillar tops are rounded, the gaps between adjacent pillars can be removed while maintaining small channels. A purely flat tile with no channels results from removing the gaps when pillar tops are flat. Therefore, four categories of texture are considered as shown in Fig. 9. To achieve reliable suction with a particle jamming substrate, urethane (Smooth-On Vytaflex30) was cast into an outer shell, filled with granular material, and clamped around an outer flange to create a seal (Fig. 8). This implementation uses two features to ensure high suction at each point of contact. First, plastic sucker back supports were fabricated to create particle free chambers into which the outer layer of urethane can collapse under vacuum, approximating the acetabular shape of an octopus sucker [26]. Second, the top surface of each suction cup was allowed to cure open to the air, with the sides wetted to part of the mold. This creates a slim meniscus at the edge of the suction cup that is not in contact with the mold itself, so that upon curing the outer surface of the infundibulum-like component of the suction structure is extremely smooth, very thin, and slightly raised \u2013 all conditions desirable for a strong suction effect upon actuation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003443_j.procir.2020.04.140-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003443_j.procir.2020.04.140-Figure3-1.png", + "caption": "Fig. 3. Vector loop model of an ideal 4-bar-mechanism", + "texts": [ + " Furthermore, the results of the mobility band serve as a basis for evaluating the influence of deviations and joint clearance on the movement behavior using a suitable sensitivity analysis technique. Thus the functionality of the mechanism can be ensured. The methodology is shown in Fig. 2. This section describes the vector loop approach for statistical tolerance analysis for a mechanism in motion. Hereby the joints are assumed to be rigid and friction within the joints is neglected whereby the computing time can thus significantly be reduced. To show the applicability of this approach, it is applied to a planar 4-bar mechanism and each of the four links is represented by a vector (Fig. 3). For the vector loop approach, the loop closure equation according to Goessner is applied to determine the degrees of freedom and thus the number of required vector loops p for a mechanism including g linkages and n joints [23]: p = g \u2212 (n \u2212 1). (1) Applying this equation to the mechanism shown in Fig. 3, which includes four linkages and four joints, it becomes apparent that one vector loop is sufficient to characterize the kinematic behavior: L1 \u00b7 e j\u00b7\u03b81 + L2 \u00b7 e j\u00b7\u03b82 \u2212 LAB \u2212 L3 \u00b7 e j\u00b7\u03b83 = 0. (2) A vector loop as shown above consists of two equations for the real and the imaginary part, which both must be equal to zero: Li \u00b7 e\u00b1 j\u03b8i = Li \u00b7 (cos\u03b8i + j \u00b7 sin\u03b8i). (3) Thus for each vector loop, two equations are provided. For the 4-bar mechanism (see Fig. 3) one vector loop consisting of two equations is sufficient to describe its kinematic behavior. In the following only the summarized form of the equation, consisting of real and imaginary part will be shown. For this non-linear vector loop equations an explicit solution is difficult to solve, so a numerical solution is required. Therefore a numerical and iterative method (e.g. Newton-Raphson or DenavitHartenberg) is applied [11]. The movement accuracy for coupling curves of relevant joints Ji j as a function of the time \u03c4 is therefore defined as the functional key characteristic (FKC): FKC(\u03c4) = Ji j(\u03c4)", + " Subsequently, this information can be used for solving the clearance vector loop equations and then be integrated into the tolerance analysis for a realistic representation of the joint clearance. Furthermore, lubricants in the non-assembly mechanism are not existing, so a contact load can occur, whereby the resulting impulse forces are transferred to following mechanical parts. These impulses and the subsequent continuous contact can be modeled by the force model [8]. According to Flores et al. [13], the size of the clearance is hereby defined by the difference in radius between bearing Rb and journal Rj: c = Rb \u2212 Rj. (5) The vector loop approach for the 4-bar mechanism of Fig. 3, is now extended by the clearance vectors c12 and c23 for the clearance affected joints J12 and J23 in Fig 6. According to Goessner (see Eq. 1), for a 4-bar mechanism consisting of four links g and four joints n, one vector loop is sufficient for the characterization of its motion behavior, whereby the informations for the clearance vectors are derived through the MBS [23]: L1 \u00b7 e j\u00b7\u03b81 + c12 \u00b7 e j\u00b7\u03b312 + L2 \u00b7 e j\u00b7\u03b82+ c23 \u00b7 e j\u00b7\u03b323 \u2212 LAB \u2212 L3 \u00b7 e j\u00b7\u03b83 = 0. (6) 3.3. Statistical tolerance analysis using sampling techniques Statistical tolerance analysis using sampling techniques offers the possibility to determine the influence of joint clearances and geometrical deviations resulting from the AM-process on the movement behavior", + " Subsequently, this information can be used for solving the clearance vector loop equations and then be integrated into the tolerance analysis for a realistic representation of the joint clearance. Furthermore, lubricants in the non-assembly mechanism are not existing, so a contact load can occur, whereby the resulting impulse forces are transferred to following mechanical parts. These impulses and the subsequent continuous contact can be modeled by the force model [8]. According to Flores et al. [13], the size of the clearance is hereby defined by the difference in radius between bearing Rb and journal Rj: c = Rb \u2212 Rj. (5) The vector loop approach for the 4-bar mechanism of Fig. 3, is now extended by the clearance vectors c12 and c23 for the clearance affected joints J12 and J23 in Fig 6. Fig. 6. Clearance vector loop model of a 4-bar-mechanism with joint clearance According to Goessner (see Eq. 1), for a 4-bar mechanism consisting of four links g and four joints n, one vector loop is sufficient for the characterization of its motion behavior, whereby the informations for the clearance vectors are derived through the MBS [23]: L1 \u00b7 e j\u00b7\u03b81 + c12 \u00b7 e j\u00b7\u03b312 + L2 \u00b7 e j\u00b7\u03b82+ c23 \u00b7 e j\u00b7\u03b323 \u2212 LAB \u2212 L3 \u00b7 e j\u00b7\u03b83 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003379_acc45564.2020.9147606-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003379_acc45564.2020.9147606-Figure1-1.png", + "caption": "Fig. 1: Chaplygin sleigh with a balanced rotor. The rotor is placed at distance of b from the rear contact.", + "texts": [ + " This simultaneous tracking is achieved by combining the calculations of matching the harmonics (the 978-1-5386-8266-1/$31.00 \u00a92020 AACC 5256 Authorized licensed use limited to: University of Wollongong. Downloaded on August 10,2020 at 06:44:26 UTC from IEEE Xplore. Restrictions apply. harmonic balance method) and a vector pursuit algorithm. In the current paper only circular and straight line reference paths in the configuration space are considered. A schematic diagram of the Chaplygin sleigh is shown in Fig. 1. The sleigh has mass m and moment of inertia I . The point P represents the point of contact of sharp knife edge or wheel with the ground. At this point the sleigh is not allowed to slip in the transverse direction. The axes Xb and Yb are body fixed where Xb is aligned with the line between P and the center of mass. The position of the center of mass of the sleigh is denoted by (x, y) and the orientation of the sleigh is \u03b8. The distance between P and the center of mass is b. The sleigh carries a balanced rotor of moment of inertia Ic at its center that is driven by a motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001769_robio.2015.7418980-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001769_robio.2015.7418980-Figure1-1.png", + "caption": "Fig. 1 Autonomous drilling robot syste", + "texts": [ + " Even so, the quality of holes still cannot be guaranteed. In recent years, industrial robot is applied in more and more factories. Production implementations of industrial robot in the aircraft manufacture and assembly are really active[2]. With the help of offline programming technology, industrial robot can be applied in more areas in factory. The autonomous drilling robot system (ADRS), mentioned in this paper, is mainly composed of an industrial robot, end effector and control system, as shown in Fig.1. ADRS can achieve high speed drilling automatically, but its accuracy is far from the requirements of aircraft assembly. Drilling perpendicularity, positioning accuracy and calibration of TCP all have great influence on ADRS accuracy[3]. The perpendicularity of the connecting holes has important effect on the quality of the riveting holes. The bad perpendicularity may lead to fatigue cracks, which could *Research supported by the National Natural Science Foundation of China (No.61375085), M.Q. Lin, P" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002258_itec.2016.7520251-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002258_itec.2016.7520251-Figure2-1.png", + "caption": "Fig 2. Scheme of an IPMSM with internal short circuits.", + "texts": [ + " By adding the flux linkage in the healthy and shorted parts, and considering that = + and \u03bc + \u03bc = 1, the total flux linkage is obtained as: = (\u03bc + \u03bc ) + (\u03bc + \u03bc ) + (2) = ( + ) \u2212 ( \u03bc ) (3) Equation (3) is important because it shows that the flux linkage of the winding is described in two components. The first component is the flux linkage expected from a healthy machine, and the second is the effect of the short-circuit given by the product of the total inductance with the shorted turns ratio and the current flowing through the failure path. Based on the previous case, the expression for the flux linkages under internal short circuits for a three-phase IPMSM as shown in Fig. 2 will be obtained. Based on [9] the terms for the magnetizing and leakage inductance matrix can be calculated as: , , ( ) = 13 + \u2212 \u2212 13 ( \u2212 ) 2 \u2212 \u2212 (4) , , = if =0 if \u2260 (5) and each of the terms of the flux linkage vector will be obtained as follow: , ( ) = ( \u2212 ) (6) where, , : a,b,c , , : Terms of the magnetizing inductance matrix Lm( ) , , : Terms of the leakage inductance matrix Lls : Rotor position , : Windings position ( = 0; = 2 /3; = 4 /3) , : Magnetizing inductance for the q and d axis : Leakage inductance for each winding : Magnet flux linkage Note that the total inductance matrix is obtained as ( ) =( ) + " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002246_978-3-319-15684-2-Figure4.5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002246_978-3-319-15684-2-Figure4.5-1.png", + "caption": "Fig. 4.5 The finite rotation angles or the robot body during walking in the down direction", + "texts": [], + "surrounding_texts": [ + "We describe an example of the turn angle ay to the left (see Fig. 4.6). We define longitudinal zi(T) and cross- xi(T) the coordinates of the center of gravity of the robot body at the time of the end of the current step T in the new rotated (i + 1)th coordinate system. The transition matrix to the (i + 1)th system of the ith coordinate system can be represented as \u00f04:1\u00de 26 V. Tereshin and A. Borina xi+1 xi 4 Control of Biped Walking Robot \u2026 27 Therefore, in the system (i + 1) transverse and longitudinal coordinates of the center of gravity can be represented as xi\u00fe1 \u00bc xi\u00f0T\u00de cos\u00f0ay\u00de ; zi\u00fe1 \u00bc 0: \u00f04:4\u00de Similarly, we write the expression for the velocity of the center of gravity in the system (i + 1), Vi\u00fe1 \u00bc Ai\u00fe1;i Vi \u00bc cos\u00f0ay\u00de 0 sin\u00f0ay\u00de 0 1 0 sin\u00f0ay\u00de 0 cos\u00f0ay\u00de 2 64 3 75 _xi\u00f0T\u00de _yi\u00f0T\u00de _zi\u00f0T\u00de 2 64 3 75 \u00bc _xi\u00f0T\u00de cos\u00f0ay\u00de _zi\u00f0T\u00de sin\u00f0ay\u00de _yi\u00f0T\u00de _xi\u00f0T\u00de sin\u00f0ay\u00de \u00fe _zi\u00f0T\u00de cos\u00f0ay\u00de 2 64 3 75 ; \u00f04:5\u00de where Ai\u00fe1;i\u2014cosine matrix, Vi\u2014column-vector of the speeds at time T onto the iaxis. Therefore, we obtain the projection of the transverse and longitudinal velocities in the rotated angle ay system i + 1 as _xi\u00fe1 \u00bc _xi\u00f0T\u00de cos\u00f0ay\u00de _zi\u00f0T\u00de sin\u00f0ay\u00de; _zi\u00fe1 \u00bc _xi\u00f0T\u00de sin\u00f0ay\u00de \u00fe _zi\u00f0T\u00de cos\u00f0ay\u00de: \u00f04:6\u00de So, rotation control is performed according to the formulas obtained in [5], adjusted for the rotation: xl \u00bc xi\u00f0T\u00de d 2 cos\u00f0ay\u00de \u00fe 1 k \u00f0 _xi\u00f0T\u00de cos\u00f0ay\u00de _zi\u00f0T\u00de sin\u00f0ay\u00de\u00de \u00f04:7\u00de where \u03b4\u2014the coefficient of stability in the transverse direction, T\u2014the end time of the current step; k \u00bc ffiffiffiffiffiffiffiffi g=L p ; L\u2014height of center of gravity; g\u2014the acceleration of free fall. We define coordinates of center of gravity of the robot body in the ith system, and let it coincide with the global [8]. The transition matrix in the global system can be represented as \u00f04:8\u00de Column-vector coordinates of the center of gravity in the global system can be represented as 28 V. Tereshin and A. Borina Ri \u00bc Hi;i\u00fe1 Ri\u00fe1 \u00bc xi\u00fe1 cos\u00f0ay\u00de \u00fe zi\u00fe1 sin\u00f0ay\u00de yi\u00fe1 xi\u00fe1 sin\u00f0ay\u00de \u00fe zi\u00fe1 cos\u00f0ay\u00de \u00fe zi\u00f0T\u00de \u00fe xi tan\u00f0ay\u00de 1 2 664 3 775: \u00f04:9\u00de Therefore xi \u00bc xi\u00fe1 cos\u00f0ay\u00de \u00fe zi\u00fe1\u00f0T\u00de sin\u00f0ay\u00de; zi \u00bc xi\u00fe1 sin\u00f0ay\u00de \u00fe zi\u00fe1 cos\u00f0ay\u00de \u00fe zi\u00f0T\u00de \u00fe xi tan\u00f0ay\u00de: \u00f04:10\u00de Figures 4.7, 4.8 and 4.9 present the results of the numerical solutions of the system of equations presented in [4], under the given conditions and parameters of the walking robot during left rotation on the second step by the angle ay \u00bc 0:3. t, c m z(t) x(t) Fig. 4.7 The longitudinal and transverse coordinates during left rotation on the second step in the global coordinate system 4 Control of Biped Walking Robot \u2026 29" + ] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure12-1.png", + "caption": "Fig. 12 Vision system of the humanoid robot iCub (Nava Rodr\u00edguez 2010): a current design with belts mechanism for the motion transmission; b proposed solution based on articulated mechanism", + "texts": [ + " This allows the hand to pass from a standby state to and operative one, ready to hold an object through different states transforming mechanisms. It is possible to see an already developed RL1 Hand in Fig. 11. This group of mechanisms make the RL1 Hand driving system to be formed by a unique Multi-State Actuator that allows the Hand to pass through different positions, since the resting states inside the anchor until the fingers and thumb driving position, necessary to hold an object. It is also possible to see those different positions. Finally, Fig. 12 shows the RL1 materializing holding tasks for different objects with different shapes and sizes. The RL2 Hand has been designed to be mounted on a new version of the MATS Robot, The ASIBOT Robot Arm that differs from the previous version by small changes at the control system, sensor system and the capacity computing resources, without changing the mechanical structure. For that reason, the predetermined requirements for the RL2 Hand are the same as the ones detailed for the RL1 Hand. The RL2 Hand was intended to enlarge the functionality of the Hand referring to shape and size of the objects to be held, basing the entire project on the experience obtained with the RL1 Hand", + " The hybrid Wheel leg of Guccione and Muscato (2003) includes a similar performance (rotational-prismatic-prismatic) although without using a pantograph. The leg mechanism of Gonzalez Rodriguez et al. (2009) also falls into the decoupled leg mechanisms but with a different structure. It has been designed in order to obtain fast and easy controlled walking operation in obstacle-free environments. A wide workspace allows the leg overcome high obstacles. The leg structure avoids joint locking and is not sensitive to reactions in any direction. According to the scheme 52 \u00c1. G. Gonz\u00e1lez Rodr\u00edguez and A. Gonz\u00e1lez Rodr\u00edguez of Fig. 12, two DOFs are required to obtain horizontal (a) and vertical (b) motion of the leg in the saggital plane. This mechanism is synthesized to provide an approximate straight trajectory segment during the support phase, only operating over one actuator, with the other being inactive. The third DOF allows rotating the plane of movement with respect to an axis parallel to the movement direction, and it is driven by a third linear actuator. It is necessary to maintain stability and to make the robot turn", + " In this case, an Mechanical Design Thinking of Control Architecture 101 interdisciplinary group of work is necessary to design the mechanical structures that contain these particular systems. The mechanical designer needs a feed-back from expert in every discipline in order to identify the possible problems that the structure can generate to the particular system, for afterward resolving them. For example, vision is a system that requires to work under high accuracy condition, in which some external effects can interfere with its proper operation. Figure 12a shows the structure of the current vision sub-system of the iCub Humanoid robot (Sandini et al. 2005). From discussions with vision experts, the following design problems have been identified in the current iCub eye sub-system (Nava Rodr\u00edguez 2010): \u2013 The belt system presents slippage and backlash since the force is transmitted by the contact between belt teeth. \u2013 In highly dynamic applications, the belt system can generate some vibrations, unsuitable for the vision system, which requires frequent adjusted of tension. \u2013 The fact that the tilt motor has to carry the pan motors slows the system down. \u2013 A small amount of backlash is always necessary to reduce excessive wear, heat and noise created by the current gearboxes. Figure 12b shows the proposed solution to the slippage, backlash and vibration problems of belt system that involves an articulated mechanism to transmit the motion from motor to eye tilt and pan. An articulated mechanism is a robust solution that provides feasible stiffness to the mechanical structure, as also reported in backlash avoidance section. The parallel manipulator that composes the eye structure of Fig. 5a can be a solution to the problem generated for the serial configuration of the eye kinematic chain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001764_978-3-319-18997-0_11-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001764_978-3-319-18997-0_11-Figure3-1.png", + "caption": "Fig. 3 Schematic diagram of simplified arc shape", + "texts": [ + " As it\u2019s known that capillary pressure is caused by surface tension, so PCdV \u00bc rLGdS PC \u00bc 2rLG R \u00f04\u00de Par is caused by arc force, which consists of electromagnetic force, plasma flow force and gas evaporation force. However, the last force is usually neglected because it is far smaller than the first two forces. Therefore, it is expressed as formula (5). Par \u00bc Fe \u00fe Fp S \u00f05\u00de It is presumed that the projection on the pool is approximate to a circle with radius Rb, Ra is regarded as the radius of tungsten electrode. S \u00bc pR2 b \u00f06\u00de The arc shape is simplified to circular truncated cone, as shown in Fig. 3. And the assumption is made that current I goes evenly through each section of the arc. Therefore, magnetic induction intensity is B B \u00bc l0H \u00bc l0i 2pri \u00f07\u00de i is the current going through the circle with radius ri. The component of electromagnetic force per unit length dl in the direction from tungsten electrode to the molten pool is df?, with the following expression: df? \u00bc sinu df \u00bc dr dl B di dl \u00bc l0ididri 2pri \u00f08\u00de The integration of df? in radial direction: dF \u00bc ZRb Ra df? \u00bc l0idi 2p ln Rb Ra \u00f09\u00de Then Fe is obtained by integrating dF: Fe \u00bc Z I 0 dF \u00bc l0I 2 4p ln Rb Ra \u00f010\u00de For TIG welding, when arc length is more than 2 mm, radial distribution of plasma flow force is in concordance with the hyperbolical curve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003929_s40430-020-02785-6-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003929_s40430-020-02785-6-Figure10-1.png", + "caption": "Fig. 10 Test rig: a Composition; b Photograph", + "texts": [ + "9\u00a0MPa, and when 11\u00b0, the MSR achieves the smallest at 42.3\u00a0MPa. In summary, when \u03b8 is 11\u00b0, the \u201cS\u201d shape structure performs the best damping effect. Three parameters of \u201cS\u201d shape structure can influence the vibration and stress in a coupled manner. The oil film clearance and distribution angle have similar effects. And the effect of the inner flange height on the vibration is opposite to the above two parameters. The vibration reduction effect verification of WDSB is conducted on a test rig, as shown in Fig.\u00a010. The motor power of the drive module is 15\u00a0kW, and the maximum rotational speed is 1200r/min. The shaft is supported by two rolling bearings, and the test bearing is suspended on the shaft. The journal at the test bearing has a ZQSn10-2 bushing with a length of 175\u00a0mm and an outer diameter of \u00d8150\u00a0mm. The loading module adopts hydraulic loading mode, the force acts on the bottom of the test bearing, and the maximum load is 1.5t. The basic parameters of CWSB and WDSB are the same (as shown in Table\u00a01), except that WDSB has a damping structure on its bushing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002468_peac.2014.7037940-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002468_peac.2014.7037940-Figure4-1.png", + "caption": "Fig. 4. Optimum stator flux trajectory derived from type A solutions for p=7", + "texts": [ + " When optimal synchronous PWM is used, the stator voltage vector of PMSM will be defined by the corresponding optimum switching angles \u03b11,\u03b12,\u2026,\u03b1N, and can be expressed as 1 2( , , , )s Nu \u03b1 \u03b1 \u03b1 . Then, the stator flux vector s\u03c8 can be derived from (6), and is expressed as 1 2( , , , )s N\u03c8 \u03b1 \u03b1 \u03b1 . For example, analyzing the type A solutions for p=7 shown in Fig. 2(c), the optimum switching angles are discontinuous. When m<0.8, the corresponding optimum stator flux trajectory in \u03b1-\u03b2 frame is derived from (6), and shown in Fig. 4(a). And the corresponding optimum stator flux trajectory when operating at 0.8\u2264m<1 is shown in Fig. 4(b). As shown in Fig. 4, the hexagonal curve of the stator flux vector will be replaced by the thirty-corner curve when the modulation index increases from m<0.8 to 0.8\u2264m<1.The small circles in Fig. 4 represent the positions where the space voltage vector is zero so that the stator flux vector s\u03c8 stops rotating and keeps unaltered. The relationships between the optimum switching angles and the optimum flux trajectory can be obtained by analyzing the Fig. 4. And the curve shown in Fig. 4(a) is chosen to discuss. As shown in this figure, the hexagonal flux trajectory are divided into 6 sectors, which are represented respectively by I, II, III, IV, V, VI. The discussion will be restricted to sector \u2160, for the trajectory has sector symmetry. In this figure, the stationary reference frame \u03b2a-\u03b2b-\u03b2c is defined, and the \u03b1-axis flux \u03b1\u03c8 and \u03b2-axis flux \u03b2\u03c8 can be resolved into \u03b2a-\u03b2b-\u03b2c components [4]. The transformation relations can be represented as 3 1 2 2 3 1 2 2 a b c \u03b2 \u03b2 \u03b2 \u03b1 \u03b2 \u03b2 \u03b1 \u03b2 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u23a7 =\u23aa \u23aa \u23aa = \u2212 \u2212\u23a8 \u23aa \u23aa = \u2212\u23aa \u23a9 (7) where \u03a8\u03b2a, \u03a8\u03b2b and \u03a8\u03b2c are \u03b2a-axis flux, \u03b2b-axis flux and \u03b2caxis flux", + " According to (8) and (9), \u03a8f1 and m are used to compute the flux thresholds and the durations of zero vectors, and the expected stator flux trajectory is set. The \u201cVoltage Reconstruct\u201d block and the \u201cFlux Observer\u201d block are used to compute the feedback flux, not introduced here, and can be realized as [8].The feedback flux \u03a8\u03b2a, \u03a8\u03b2b and \u03a8\u03b2c, the phase of the feedback flux fb\u03b3 ,\u03b3 and the flux thresholds are all used to generate the output pulses in the block \u201cPulse Generation\u201d, where the difference between fb\u03b3 and\u03b3 are used to end the zero vector. In addition, as shown in Fig.4, the switching between the different PWM patterns means the transition between the different flux trajectories. To maintain a continuous stator flux fundamental and reduce the peak value of the current, the switching should be done near the end point of the current sector. The proposed closed-loop synchronous PWM has some merits. a) Under steady-state conditions, the harmonic distortion of the motor currents is minimized. b) Under dynamic operation conditions, the output pulses are dynamically controlled by regulating the stator flux thresholds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003308_s40430-020-02488-y-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003308_s40430-020-02488-y-Figure14-1.png", + "caption": "Fig. 14 Spatial meshing coordinate system", + "texts": [ + " However, their research has not been simulated or tested, and the concrete effect remains to be verified. The tooth surface equation for hobbing can be written in the form of: In Eq.\u00a0(9), 1 is the angle at which the hob turns, 2 is the angle at which the gear turns, is the distance between the hob spindle and the gear blank spindle, is the angle between the hob spindle and the gear blank spindle, and x1 , y1 and z1 are the coordinates of the hob surface, which can be obtained by Eq.\u00a0(3), as shown in Fig.\u00a014. The development of CAD/CAM technology makes mathematical models and simulation analysis indispensable in the design and manufacture of LSBG. This paper reviews the development history of logarithmic spiral bevel gears and summarizes the key principles and methods of this new gear from theory to practical processing applications, including the establishment of LSBG mathematical model, finite element simulation analysis, processing and manufacturing methods. These laws and methods are universal. If other researchers study a new type of gear, they can learn from the principles and methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001463_iciea.2015.7334434-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001463_iciea.2015.7334434-Figure4-1.png", + "caption": "Fig. 4. The relationship between voltage vector and stator flux vector.", + "texts": [ + " Finally, according to space position of stator flux, choose appropriate voltage space vectors in switch table, realize motor control. Neglecting stator resistance, relationship between stator flux and voltage is expressed as (11). s s s s s d dR dt dt = + \u2248u i (11) In transient process of motor, because sampling time is very short, increment of stator flux s\u03c8\u0394 can be gotten by calculating product of stator voltage vector su and t\u0394 , and direction of s\u03c8\u0394 is same as stator voltage vector su . The relationship between voltage vector and stator flux vector is shown in Fig. 4. According to Fig.4, stator flux vector s can be written as: s s sje \u03b8\u03c8= (12) Where \u03b8s is the angle between s and \u03b1 axis (degree). Combine (11) with (12), there is: s s s s sr sn sjd e j u ju dt \u03b8\u03c8 \u03c9= + = +u (13) Where \u03c9s is angular velocity of stator flux (rad/s), sru is radial component of su (V) and snu is normal component (V). According to (13), stator flux vector s can be controlled by suitable voltage vector su directly, and radial component sru controls amplitude of flux, normal component snu controls angular velocity of flux, load angle sm\u03b4 can be expressed as: sm s r( )dt\u03b4 \u03c9 \u03c9= \u2212 (14) 1966 2015 IEEE 10th Conference on Industrial Electronics and Applications (ICIEA) When direction of sru of given voltage vector is same as stator flux s , amplitude of s is increased, on the contrary, it is decreased" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003272_j.promfg.2020.05.027-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003272_j.promfg.2020.05.027-Figure1-1.png", + "caption": "Figure 1. Traditional longitudinal-excited vibrator design to achieve longitudinal and torsional vibrations: (a) diagonal slits type [2\u20134,17]; (b) helical grooves type [11,12].", + "texts": [ + " To optimize the design parameters, a finite element (FE) analysis was carried out in studying the vibration modes. Laboratory experiments were conducted to validate the electromechanical and vibration performance of the new design of L&T ultrasonic vibrator. Details are discussed in the following sections. There are two conventional designs to deliver longitudinal vibration to L&T vibration by modifying external structures. One is adding a series of diagonal slits around the circumference of the vibrator, as shown in Figure 1(a). In this case, the axially propagating wave can be reflected in the slits and produce the desired L&T hybrid vibration at the end of the vibrator [3,4,17]. The composited vibration is affected by the size, depth, width, and length of the slits. Another method is to use helical grooves on the vibrators [11,12], as shown in Figure 1(b). Compared to the first method of using straight slits arrays, the helical grooves are more difficult to be machined due to the complex geometric shapes, even though the torsional conversion efficiency sometimes can be higher than that of the diagonal slits. However, both the two methods require the resonant frequency of the torsional vibration mode to be the same as that of their longitudinal mode. This is challenging particularly when they are sharing only one node plane. Inspired by the idea of modifying the structure of the longitudinal-excited vibrator, this paper proposed to design a waveguide structure with internal and external geometric features" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure20-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure20-1.png", + "caption": "Figure 20. Experimental setup to test the gears with conical and flat passive rollers as worm wheels", + "texts": [ + " Even though the dynamic frictional resistance increased in the experiments of the spur and crowning gears, the dynamic frictional resistance of the gear with flat passive roller was almost constant and as small as the value of \u201cFree slide\u201d, hence confirming the advantage of the gear with flat passive rollers as compared to dynamic frictional resistance. In a free sliding state, the omnidirectional gear didn\u2019t touch any other gears, regardless of the kind of gears on the omnidirectional gear. Accordingly, even though this experiment was performed under the same condition as Fig. 15, we plotted a horizontal line to show free sliding only in Fig. 19, to show a typical and common example effectively. In this research, we also examined the advantage of the gear structure with passive rollers as a worm wheel shown in Fig. 20 to transmit power from a worm gear to the other direction perpendicular to its own. In the case of the ordinary worm wheel, power transmission efficiency tends to be low because of the high frictional resistance in sliding motion between the teeth of the worm gear and the worm wheel [10-12]. Limiting the rolling resistance of passive rollers on the developed gear may solve this problem. We used the experimental setup shown in Fig. 20 to compare the power transmission efficiencies of the ordinary worm wheel and gears with conical and flat passive rollers, using pulleys with the same diameter on the input and output shafts. The force gauge with Kevlar wire pulls one side of the pulley to input torque on the worm gear. This torque is then transmitted to the worm wheel, and the enhanced torque is measured by the other fixed force gauge on the Kevlar wire of the output pulley. The teeth width of the first ordinary worm wheel (helical wheel) for comparison is 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000054_iccia49288.2019.9030886-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000054_iccia49288.2019.9030886-Figure2-1.png", + "caption": "Figure 2. Kinematic schematics of a deployable suspended cable driven robot.", + "texts": [ + " Next, actuators and power transmission dynamics are formulated and the overall dynamics of the robot is derived. Then, the proposed robust controller is introduced and its robust stability is analyzed. Finally, the experimental results are presented and the suitable performance of the controller is shown. This section is devoted to deriving the kinematics and dynamics formulation of the DSCR. To this end, first the inverse kinematics of the robot is derived, next. Finally, the dynamics formulation of the robot including its actuators is presented. Fig. 2 illustrates a DSCR with four cables. As shown this figure, all the cables are attached to a single point on the end-effector. Typically, the end-effector is modeled as a lumped mass located at the point of cable intersections. The loop closure for this manipulator is formulated as follows [21]. \u2212\u2192 Li = \u2212\u2192 P \u2212\u2212\u2192 PAi , i = 1, . . . ,4. (1) Algebraically rewriting the loop closure equations yeilds: (li)2 = (P\u2212PAi) T (P\u2212PAi), (2) where li is the length of i\u2019th cable. component wisely expanding this equation leads to: li = \u221a (x\u2212 xi)2 +(y\u2212 yi)2 +(z\u2212 zi)2 (3) in which, the x,y,z and xi,yi,zi, denote the position of endeffector and attachment points of the i\u2019th cables, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001333_ilt-10-2011-0080-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001333_ilt-10-2011-0080-Figure3-1.png", + "caption": "Figure 3 Simplified geometric model", + "texts": [ + " (2003) investigated oil-film formation under reciprocating rolling point contact experimentally and numerically. Wang et al. (2005a, b) numerically studied the TEHL line contact problem under reciprocating motion between two rollers and then further observed the point contact EHL films under pure rolling short stroke reciprocating motion theoretically and experimentally. In the previous investigation of TSUJ, Wang and Chang (2009b, 2010) and Chang et al. (2010) observed its lubricating performance by using the simplified geometric model (Wang and Chang, 2010) shown in Figure 3 under isothermal Newtonian flow conditions and then further studied the effect of the effective radius on the lubricating behavior thereof considering thermal effect. In this work, the lubricating properties of the TSUJ were also investigated by using the simplified geometric model shown in Figure 3 under thermal Newtonian flow condition. The effects of the amplitude and frequency of the reciprocating motion on the pressure, the film thickness and the temperature were discussed. According to Figure 3, the relative motion between the plane a and the roller b is the simple reciprocating sliding motion. Assuming the plane a kinetic whereas the roller b static, the velocity of the plane a can be defined as: u \u00bc Av sinv \u00b7 t \u00f01\u00de where v \u00bc 2pf, and symbols A, f and t stand for the amplitude (1/2 of a stroke), the frequency and the time value of reciprocating motion, respectively. Effects of reciprocating motion parameters X.F. Wang, R.F. Hu, K.S. Wang and H.R. Cui Volume 66 \u00b7 Number 1 \u00b7 2014 \u00b7 111\u2013123 According to Yang and Wen (1990), Reynolds equation under thermal Newtonian flow model can be simplified as: \u203a \u203ax \u00f0r=h\u00deeh 3 \u203ap \u203ax \u00bc 6u \u203a\u00f0 ~rah\u00de \u203ax \u00fe 12 \u203a\u00f0reh\u00de \u203at \u00f02\u00de where: \u00f0r=h\u00dee \u00bc 12\u00f0her 0 e=h 0 e 2 r00e\u00de ~ra \u00bc 2\u00f0re 2 r0ehe\u00de r0e \u00bc \u00f01=h2\u00de Z h 0 r Z z 0 1=hdz0dz re \u00bc \u00f01=h\u00de Z h 0 rdz r00e \u00bc \u00f01=h3\u00de Z h 0 r Z z 0 z0=hdz0dz he \u00bc h= Z h 0 1=hdz h0 e \u00bc h2= Z h 0 z=hdz where symbols r h, h and p represent the density, the film thickness, the viscosity and the pressure of lubricant, respectively; symbols x and z symbolize horizontal and vertical coordinates, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002641_cgncc.2016.7829027-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002641_cgncc.2016.7829027-Figure1-1.png", + "caption": "Figure 1. Engine distribution of heavy lift launch vehicle", + "texts": [ + " So this paper designs a new adaptive augmented fault tolerant control algorithm, which includes four parts: basic PID controller, fast on-line fault detection and reconstruction controller, adaptive vibration frequency identification algorithm and adaptive gain adjustment law to deal with variety of issues such as multiple interference patterns, inaccurate vibration frequency measurement, proneness to failure of executive bodies in heavy lift launch vehicle. During ascent flight phase, heavy lift launch vehicle is mainly affected by external disturbance forces such as engine thrust, aerodynamics, gravity, Coriolis force, inertial force of the engine swing, generalized force due to vibration, etc. The coordinate system definition and coordinate transformation relation during active segment flight are referenced in [13]. The engine distribution of heavy lift launch vehicle proposed in this paper is shown in Fig.1, with 4 core stage engine two -way swing and 4 booster stage engine tangential swing (8 engine specifications are consistent). With a total of 12 engines, system redundancy can be improved and flight mission failure can be avoided in the case of engine failure. 978-1-4673-8318-9/16/$31.00\u00a92016 IEEE As shown in Fig.1, 12 engine swing angle corresponding to 8 engines can be converted to equivalent three-channel independent engine swing angle \u03d5 \u03c8 \u03b3\u03b4 \u03b4 \u03b4 . From 2E Cr r= , it can be obtained: 1 3 5 7 10 12 2 4 6 8 9 11 2 3 6 7 9 10 11 12 6 6 2 2 2 2 8 \u03d5 \u03c8 \u03b3 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4\u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4\u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4\u03b4 + + + + + = + + + + + = \u2212 \u2212 + + \u2212 \u2212 + + = (1) For accurate and complete description of movement of heavy lift launch vehicle, centroid dynamics of heavy lift launch vehicle can be established under launching coordinate system, and rotation dynamics equation around the center of mass can be established under launch vehicle coordinate: 2 ' 2 d ( ) 2 d d d c k e e e T S D B m m m m t t = + + + + \u2212 \u00d7 \u00d7 \u2212 \u00d7 + \u00d7 = + + + + r P R F g F r H H M M M M M (2) where, r represents position vector of the rocket at launching coordinate system, P represents engine thrust, R is aerodynamic force, cF is control force, mg is gravity term, ' kF is additional relative force, ( )e em \u00d7 \u00d7 r is centrifugal inertial force, 2 em \u00d7 is Coriolis inertial force, is launch vehicle speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003166_0957456520923319-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003166_0957456520923319-Figure9-1.png", + "caption": "Figure 9. First modal natural frequency of composite specimen.", + "texts": [ + " The three-point bend test is carried out in the glass fiber\u2013reinforced composite specimen to determine and validate the modulus value obtained from the tap test. The modulus from three-point bend test is 15,467.13 MPa, and it is observed that the moduli obtained from both the tests are in good agreement. The obtained mechanical property of the C-glass fiber composite specimen is inserted in the engineering data, and the modal analysis (using ANSYS) is carried out to determine the natural frequency of the specimen. From the tap test, the natural frequencies of the specimen are 28.9 and 29.143 Hz from the analytical method. Figure 9 shows the first modal natural frequency of the composite specimen. The error percentage between the experimental and numerical values is 0.84%, which indicates that the natural frequencies obtained from both the tap test and ANSYS have good accuracy. A three-point bend test is carried out to determine the load versus deformation curve for steel master leaf and composite graduated leaf. Figure 10 shows the deformation versus load curve, and it is observed that at the 1600 N load, deformation in steel spring is 114" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001296_01691864.2014.959051-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001296_01691864.2014.959051-Figure7-1.png", + "caption": "Figure 7. The analysis model of Climbot when transiting.", + "texts": [ + " To answer this question, the additional constraints on the fiveDoF Climbot are addressed in this subsection, leading to the D ow nl oa de d by [ M cM as te r U ni ve rs ity ] at 1 3: 06 0 9 Ja nu ar y 20 15 Algorithm 3 The binary approximating method 1: Construct the workspace sphere 2: Calculate the intersections of the workspace sphere and the axis of the target pole 3: if no intersection exists then 4: return G R \u21d0 \u2205 5: else 6: Validation tests with P A, P B and PC , encoding ABC 7: if ABC = 000 then 8: for each hal f part do 9: Call the valid mid-point searching function, returning bL and bR 10: if bL = \u2205 then 11: Call the binary approximating function, returning G Rhal f 12: else 13: G Rhal f \u21d0 \u2205 14: end if 15: end for 16: return Union set of the G Rhal f from each half part 17: else if ABC = 010 then 18: Call the binary approximating function, returning G R 19: return G R 20: \u00b7 \u00b7 \u00b7 / \u2217 Processing other cases listed in Table 1 \u2217 / 21: end if 22: end if approach to figure out adaptive potential configurations for later validation in the binary approximating method. Note that all configuration parameters (vectors and matrixes) in this subsection are described with respect to the base frame {B}, so the leading superscript B is omitted for simplicity of the expression. Due to the special mechanical configuration, the links of Climbot are always in a plane (which we call the \u2018robot plane\u2019hereafter, painted with gray in Figure 7). Suppose the target-grasping position and orientation for the transition configuration are Pg = [ pgx pgy pgz ]T and GR = [n o a], respectively. The following constraints must be satisfied, k \u00b7 a = 0, n \u00b7 a = 0, \u2016a\u2016 = 1, (11) where k = [\u2212 sin \u03b1 cos \u03b1 0]T represents the norm vector of the robot plane, and tan \u03b1 = ay/ax = pgy/pgx , as shown in Figure 7. Equation (11) associates the orientation with the position from the perspective of the grasping feasibility with fiveDoF serial robots like Climbot. If the grasping orientation GR is known (for example, to grip square poles), the grasping position can be computed directly as,{ t = p0x ay\u2212p0yax nyax \u2212nx ay Pg = P0 + t \u00b7 Pdir . (12) If the grasping position Pg is given (for example, to grip a cylindrical pole at a specific point), the grasping orientation can be also computed as, \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 n = Pdir a = [ \u00b11/ \u221a 1 + tan2 \u03b1 + \u03bb2 ax tan \u03b1 \u03bbax ]T o = n \u00d7 a , (13) where \u03bb = \u2212(nx + ny tan \u03b1)/nz " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000969_icuas.2013.6564700-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000969_icuas.2013.6564700-Figure1-1.png", + "caption": "Fig. 1. Diagram of a UAV at position ~rA that is moving at heading angle \u03b8 and tracking a randomly-moving target at ~rT with distance r = |~rA \u2212 ~rT | and relative angle \u03d5.", + "texts": [ + " Section III reviews the methodologies for accurate discretization and value iteration to compute the control. In Section IV we describe the resulting control law and demonstrate the effectiveness of this approach when the tracker can lose sight of its target. Section V concludes this paper and provides directions for future research. We consider a small, fixed-wing UAV flying at a constant altitude in the vicinity of a groundbased target, tasked with maintaining a nominal distance from the target. The target is located at position ~rT (t) = [xT (t), yT (t)] T at the time point t (see Fig. 1). The UAV, located at position ~rA(t) = [xA(t), yA(t)] T , moves in the direction of its heading angle \u03b8 at a constant speed vA. In our problem formulation, the target motion is unknown. We therefore assume that it is random and described by a 2D stochastic process. Drawing from the field of estimation, the simplest signal that can be used to describe an unknown trajectory suggests that the motion of the target should be described by 2D Brownian motion: dxT (t) = \u03c3dwx (1) dyT (t) = \u03c3dwy (2) where dwx and dwy are increments of unit intensity Wiener processes along the x and y axes, respectively, which are mutually independent", + " With the choice of \u03b8c = \u03b8 + u/\u03b1\u03b8, the UAV can be modeled as a planar Dubins vehicle [2]: dxA(t) = vA cos (\u03b8(t)) dt dyA(t) = vA sin (\u03b8(t)) dt d\u03b8(t) = udt, u \u2208 U (6) where the turning rate u \u2208 U \u2261 {u : |u| \u2264 umax} must be found. In order for the control to be independent of the heading angle of the UAV or the absolute position of the UAV or target, we relate the problem to relative dynamics based on a time-varying coordinate system aligned with the direction of the UAV velocity. The reduced system state is composed of the distance between the UAV and target r = |~rT \u2212 ~rA| and the viewing angle \u03d5 between the UAV\u2019s direction of motion and the vector from the UAV to the target, as seen in Fig. 1: r = \u221a (\u2206x)2 + (\u2206y)2 (7) \u03d5 = tan\u22121 ( \u2206y \u2206x ) . (8) The combined UAV-target system should maintain the relative distance r at the nominal distance d for all times. To this end we seek to minimize the expectation of an infinite-horizon cost function V (\u00b7) with a discounting factor \u03b2 > 0: V (r, \u03d5, \u03c4) = min u\u2208U E {\u222b \u221e 0 e\u2212\u03b2t(r(t)\u2212 d)21{\u03c4=0}dt } . (9) A high value of \u03b2 places more weight on the instantaneous cost, while \u03b2 near 0 considers future costs, as well as instantaneous cost. In order to avoid any claims of certainty equivalence, the UAV should decide its control based only on its observations of the relative target state at \u03c4 = 0 and not on predictions of the target position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003791_tase.2020.3035438-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003791_tase.2020.3035438-Figure3-1.png", + "caption": "Fig. 3. Introduction of multidimensional force collection with M3713B.", + "texts": [ + " In the structure of the system, as shown in Fig. 2, the trajectory data collected by the camera-based motion analyzer are smoothed by a Gaussian smoother. The smoothed trajectory is mapped into the PTSR workspace with an alignment of PTSR\u2019s initial position. The tested HDP and the end-effector of PTSR were connected rigidly with a sixaxis force sensor (M3713B, Sunrise Instruments Co., Ltd.). The force sensor could monitor the interface force/torque between the tested HDP and the PTSR. The direction of six forces/torques was presented in Fig. 3. A pressure measuring treadmill (FDM system, zebris Medical Gmbh, German) was used to set constant speed and measured the GRF of the prostheses. Finally, the final joints\u2019 information of PTSR is reordered by the ANN controller to improve the stability of the end-effector. As the lengths of the whole measured prosthetic legs (including the thigh, shank, and foot), were not the same, a column elevator was used to adjust the PTSR\u2019s basement height. The gait parameters studied in this research could be divided into three categories: kinematics, kinetics, and mutual information [12]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002135_amc.2016.7496392-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002135_amc.2016.7496392-Figure8-1.png", + "caption": "Fig. 8. Structure of arm and leg of the robot", + "texts": [ + " They are also the oposite lines of outside circular arc lines of the movilable regions in Fig.6. The enclosed region by the four lines represents the COM Reachiable Region(CRR). Generally, the shape of humanoid robot is similar to real human. But most of all humanoid robots have more small number of degrees of freedom, although the human have the seven degrees of freedom in the arm from the shoulder to the wrist , and the six degrees of freedom in the leg from the hip joint to the ankle. Therefore, humanoid robot cannot always perform the same pose of real human. Fig.8 shows an example of the arm and leg of a prototype of humanoid robot. The arm consists of four joints and the leg consists of six joints. Therefore, there is the pose nonavailable to perform relate to human. The solvability is depend on wether the inverse kinematics problem can be solved or not. Now, it is assumed that the four limbs are located on their corresponding holds and the balance is kept. The Fig..9 expresses such the situation. This figure illustrates the sequence via three poses expressing the robot climbing from the lower position to the upper position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001950_gt2013-95086-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001950_gt2013-95086-Figure2-1.png", + "caption": "FIGURE 2. SCHEMATIC REPRESENTATION OF TAPERED VBD BRUSH SEAL.", + "texts": [ + " Although the idea of using thick bristles near the backplate to withstand pressure loading, as applied to the multi-layered brush seal (Patent # US5480165) is retained, the new feature in the VBD brush seal is the fine bristles sandwiched between a layer of thicker bristles and a layer of thickest bristles so that the fine bristles are protected from wear [4] and flow-induced vibration and high bending stress, while the fine bristles, as the core of the seal, do the sealing work and provide flexibility. In #US52101530A a seal with a tapered bore is presented, while the seal described in Patent #US5480165 utilized a stepped bore. In another embodiment of the VBD brush seal, the bristle tip can be tapered to reduce the impact of thickest portion on heat generation and rotordynamics (Figure 2) In this paper, we present the results of experimental characterization of VBD brush seal leakage, stiffness and wear. The results are presented for four different designs of VBD brush seals, as shown in Table 1. The test data presented in this paper has been normalized in order to allow focus on relevant physical phenomena instead of proprietary details. Force measurements (Figures 4,5 and 8- 11) have been normalized by the maximum force measured during the zero pressure loading of Design #4 (see Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001676_amm.823.161-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001676_amm.823.161-Figure9-1.png", + "caption": "Fig. 9. The biomechanics system of the hip joint prosthetic divided into finite elements", + "texts": [ + " To have a correctly entering of the stem prosthesis, the intramedullary canal was virtual widened, assuming a similar shape to that of the femoral stem (Fig. 6). In a sharp pelvic bone was performed similar spherical outer surface as the polyethylene cup (Fig. 7). Finally, using movement and simple geometric constraints, but based on orthopedic information (sagittal angle of 45\u00b0 and anteversion angle of 12\u00b0 for the polyethylene cup) was obtained prosthetic hip joint model, shown in Figure 7. The dividing operation into finite elements for analyzed biomechanical system was performed under the same conditions (Fig. 9) [4,5,6]. To obtain the mesh structure, the finite elements of tetrahedron type have been used and their network distribution for entire biomechanical prosthetic hip model contains a total of 128,446 elements and 23,125 nodes. This model can be used to determine the behavior of biomechanical system components in different situations. All results will be suposed to critical analysis. For pelvic bone and for femur bone the material properties were assumed to be homogeneous. Young's modulus and Poisson's ratio for trabecular and cortical bone were taken to be 70 MPa and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002900_j.engstruct.2020.110218-FigureA.13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002900_j.engstruct.2020.110218-FigureA.13-1.png", + "caption": "Fig. A.13. Force analysis of fold-line experiments with apparatus gravities considered.", + "texts": [ + " Gravity from the testing apparatus itself may reduce the accuracy of fold-line strength measurement, especially in thin-plate fold-lines. As discussed in Section 2.1.2, =M M Asinoffset max is the moment reduction caused by the gravity weight of the steel arm, where A is the rotation angle, Mmax is the maximum gravity moment at 90\u00b0 rotation. Mmax is 0.17445 Nm for the current testing apparatus. To calculate the Moffset , we separate the gravity into two major parts, the square part of the loading arm (G1) and the bar part of the loading arm (G2), as shown in Fig. A.13. The detailed calculation ofMoffset is given in Eq. (7), where h1 is the arm length ofG1, h2 is the arm length ofG2, h1 ' is the maximum arm length ofG1, h2 ' (equal to h3) is the maximum arm length of G2. G1, G2, h1 ' and h2 ' are 2.55 N, 1.61 N, 45 mm and 40 mm, respectively. Eq. (8) gives the relationship among the measured moment Mmeasured, the true hinge moment Mhinge and the Moffset , where F is the fore applied by the load cell, h3 is a constant force arm of F . = = + = +M M A G h G h G h G h Asin ( )sinoffset max 1 1 2 2 1 1 ' 2 2 ' (7) = =M M M Fhmeasured hinge offset 3 (8) The gravity from the apparatus only significantly affects the response curves of 1 mm fold-lines in this study, because the gravity correction is less than 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002298_101686616x14555429843762-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002298_101686616x14555429843762-Figure6-1.png", + "caption": "Fig. 6: Geometric objects of the FE model and meshing.", + "texts": [ + " Now adding the temperature-induced wedge imperfection to that of the allowable misalignment (as per Ref. [8] \u00a77.4) gives 1/1500 + 1/1000, yielding 1/600. This is a wedge imperfection magnitude that should not be neglected in the design of a roller bearing. The next two sections are dedicated to studying the effects of applying this imperfection to traditionally designed roller bearings. For the FE analyses, a commercial code was used. The geometry of the FE model of the roller bearing is depicted in Fig. 6a and b, and shows the roller cylinder (blue), the four contact plates (red), the support plates (green) as well as part of the bridge girder (orange) and the abutment (gray). Each roller consists of two roller drums separated by a reduced diameter part in the middle; the sideways motion control rib is accommodated here (yellow). The rotary motion of the roller is controlled by tooth-edged wheels and corresponding rulers at either end (visible in Fig. 4), which are not included in the FE model. 210 Scientific Paper Structural Engineering International Nr", + " Isotropic hardening von-Mises plasticity was assumed for both materials. The upper support plate is attached to the flange of the girder and the lower support plate is lying on a mass of concrete: full mechanical locking was assumed in these two interfaces. The analyses presented hereafter were performed for a stiffener position directly above the roller. All parts of the assembly were modelled using the element type C3D8R, which is a continuum eightnode, first order, reduced-integration volume element. Fig. 6c shows the FE mesh of the roller assembly. To study the nature of the problem a very fine mesh is required at the contact surfaces of the roller so that the distribution of the contact stresses is accurately captured. In fact 20 nodes across the width of the contact zone were included in the model, which are of the order of 3 to 4 mm. To this end, separate geometrical sub-objects, where a much finer mesh is generated, are defined along the contact lines, i.e. at the top and bottom regions of the drums as well as at the corresponding areas of the contact plates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000919_vppc.2014.7007108-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000919_vppc.2014.7007108-Figure1-1.png", + "caption": "Fig. 1. A clutch cover assembled [4].", + "texts": [ + " Therefore, the aim of this work is to propose and implement a new methodology, based on multivariate statistical analysis, to efficiently predict judder behavior through the torque signal of different facing materials obtained on a real bench test. II. CLUTCH SYSTEM The main function of a clutch system is to engage and disengage the engine to the gearbox for shifting gears, with most comfort for the drivers as possible. Basically, a clutch system can be divided into three parts: clutch cover assembled, disc and releaser system. Figure 1 illustrates a clutch cover with its main parts highlighted. The clutch cover applies the normal force on the disc by the diaphragm spring. The disc is connected to the gearbox by the input shaft and the torque T is transmitted when the clutch is engaged, calculated by the following equation [3]: T F i r\u03bc= \u00d7 \u00d7 \u00d7 (1) where F is the normal force (diaphragm spring), \u03bc the coefficient friction, i the number of surfaces contact and r the average ratio of the facing. The diaphragm spring is responsible for the clamp load and the pressure plate for applying this force on the disc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002755_1.5063139-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002755_1.5063139-Figure6-1.png", + "caption": "Figure 6: BC", + "texts": [ + " To automate the SLM process simulation the following custom scripts have been developed:: 1. Create a new transient thermal simulation module in ANSYS (V15) workbench 2. Import a material properties (XML file) 3. Import geometry (STP, IGES, ACIS, CATIA) 4. Determine along which direction the part will be built. i.e. platen normal 5. Slice part into concatenated layers of required thickness 6. Apply Symmetry if possible 7. Mesh the part 8. Determine each layers position in space relative to the build platform 9. Identify upper layer surface(s) (Figure 6) 10. Based on custom algorithms: 10a. Insert new heat flux every load step 10b. Delete previous heat flux 10c. Kill material not yet in the simulation 10d. Bring to life material each load step 10e. Set solution settings (e.g. substeps) 11. Apply temperature ( 200\u00b0C) to the platten 12. Run the simulation Steps 4-5 were completed using a script within ANSYS Design modeller. This was done due to the repetitive nature of creating multiple slices via body operation. Steps 8-10 were completed using custom algorithms inserted via Javascript code", + "2 mparison of e and 500\u00b5 dual lattice possible to s simplificati ly 75% of concatenati approximat in compar C with double f part failu iled comple diameters a yer of the he simulate dy compon ate the s me of 10 se W/mm2. lement densit m (bottom) cubes are s simulate ju on results in required ele on and dou ely 98% of ison to the symmetry co re tely in the c nd melting top \u2018cap\u2019 ( d temperatu ents, it is po imulation. conds was y at 50\u00b5m (top ymmetric in st one-quart a reduction ments. Wh ble symme elements c full simulati ndition applied ase study h occurred fro Figure 6). re results w ssible to ver A simulat applied with ) 2 er of en try an on ad m By ith ify ed a Fr co of to S m A te fa om this sim mponent te the materia have failed LM process aterial prope Table 3: M Load ste 0 +1 +2 +3 +4 +5 +6 +7 s can be mperature a ilure as mor ulation it w mperature ex l by more th . This does exactly, but rties sufficie aximum Temp p RAD0.5 354.92 425.87 466.02 499.5 528.98 555.58 579.95 597.02 seen in Ta t the first c e layers of m ximum temper first cap laye ded (0). Blue \u2013 Green r1 as observe ceeds the m an 5\u00b0C it is not reflec it does emu ntly to predi erature at star RAD1 307" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000340_jae-141878-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000340_jae-141878-Figure11-1.png", + "caption": "Fig. 11. (a) Experimental comparison of deflection of cage rotor with inner rotor for a 76.2 mm stack length, (b) Diagram showing how the cage bar sits above the inner rotor. The cage rotor bar is fixed on either ends and deflection occurs over the 90 mm of length. However, the active region is 76.2 mm.", + "texts": [ + " A summary of the deflection results for larger axial lengths is shown in Table 3. The deflection analysis presented above has been validated by experimentally measuring the deflection when using an MDL400 Tesa-Hite height gauge for the case when only the inner rotor and one cage rotor bar is present. Figure 9 shows the inner and outer rotor of the FFMG with 76.2 mm active stack length. The deflection obtained from FEA is shown in Fig. 10. The comparison with experimental deflection is shown in Fig. 11. The deflection accuracy is not as high as for the case when using a 6 inch axial length [6] however the results are sufficiently accurate for design purposes. In this paper it has been shown that the amount of radial deflection on the magnetic gear cage rotor bars for both the scaled-up and subscale magnetic gear increases linearly with axial length. The level of deflection is not significantly changed by the cage rotor diameter. The paper also demonstrates that, for the axial lengths considered, the flux-focusing magnetic gear torque density is not significantly affected by the changes in the axial length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.31-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.31-1.png", + "caption": "Figure 9.31 Special poses of the manipulator with c\ud835\udf035 = 0.", + "texts": [ + "469) c13 = ut 1eu\u03033\ud835\udf191 eu\u03032\ud835\udf192 eu\u03033\ud835\udf193 u3 = ut 1eu\u03033\ud835\udf191 eu\u03032\ud835\udf192 u3 = c\ud835\udf191s\ud835\udf192 (9.470) c23 = ut 2eu\u03033\ud835\udf191 eu\u03032\ud835\udf192 eu\u03033\ud835\udf193 u3 = ut 2eu\u03033\ud835\udf191 eu\u03032\ud835\udf192 u3 = s\ud835\udf191s\ud835\udf192 (9.471) Afterwards, with the availability of \ud835\udf031, the other joint variables can readily be obtained from Eqs. (9.450), (9.451), (9.452), (9.457), and (9.459). \u2217 Special Solution with c\ud835\udf3d5 = 0 As indicated above concerning Eqs. (9.458) and (9.459), the general solution procedure cannot be used if it turns out that c\ud835\udf035 = 0 or \ud835\udf035 = \ud835\udf0e\u20325\ud835\udf0b\u22152 with \ud835\udf0e\u20325 = \u00b11. Such special poses of the manipulator are illustrated locally in Figure 9.31. In such a special pose, which resembles the third kind of position singularity of the elbow manipulator, the solution for the other joint variables can be obtained as explained below. With \ud835\udf035 = \ud835\udf0e\u20325\ud835\udf0b\u22152, Eq. (9.436) can be written as follows: C\u0302 = eu\u03033\ud835\udf031 eu\u03032\ud835\udf033 eu\u03031\ud835\udf0e \u2032 5\ud835\udf0b\u22152e\u2212u\u03033\ud835\udf03 \u2032 6 eu\u03031\ud835\udf0b = eu\u03033\ud835\udf031 eu\u03032\ud835\udf033 eu\u03032\ud835\udf0e \u2032 5\ud835\udf03 \u2032 6 eu\u03031\ud835\udf0e \u2032 5\ud835\udf0b\u22152eu\u03031\ud835\udf0b \u21d2 eu\u03033\ud835\udf031 eu\u03032(\ud835\udf033+\ud835\udf0e\u2032 5\ud835\udf03 \u2032 6) = C\u0302e\u2212u\u03031\ud835\udf0be\u2212u\u03031\ud835\udf0e \u2032 5\ud835\udf0b\u22152 = C\u0302e\u2212u\u03031\ud835\udf0e \u2032 5\ud835\udf0b\u22152e\u2212u\u03031\ud835\udf0b (9.472) The variables \ud835\udf033 and \ud835\udf03\u20326 cannot be found separately from Eq. (9.472) because they happen to be rotation angles about parallel axes, which are the axes of the third and sixth joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002077_10402004.2016.1177151-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002077_10402004.2016.1177151-Figure1-1.png", + "caption": "Fig. 1 Flexure pivot tilting pad gas bearing with metal mesh dampers in parallel", + "texts": [ + " However, low damping remains the main barrier in gas-lubricated bearing systems because of the inherently low viscosity of gas. Low damping in the rotor-bearing system leads to large rotor vibrations during critical speed transitions and subsynchronous excitation of rotor eigenvalues at high speed. Thus, two metal mesh dampers (MMDs) are inserted into each side of the bearing housing to improve the effective damping of flexure pivot tilting pad gas bearings with pad radial compliance (FPTPGB-Cs) in Ref. (15). Fig.1 shows the structure of FPTPGB-Cs with MMDs in parallel (FPTPGB-C-MMDs). Two MMDs are inserted into each side of the bearing housing parallel to the pad. The bearing can be simply considered as a gas film in series with the bearing support. As the ratio between gas film stiffness and bearing support stiffness increases, the equivalent damping of the bearing system approaches the damping of the bearing support (15). Thus, the effective damping of the bearing is improved because of the superior Coulomb damping mechanism of MMDs (16). Thus, the D ow nl oa de d by [ U ni ve rs ity o f N eb ra sk a, L in co ln ] at 1 1: 08 0 6 Ju ne 2 01 6 ACCEPTED MANUSCRIPT rotor system supported by FPTPGB-C-MMDs can stably operate at high speed, thereby improving the stability and power density of the turbomachinery. [insert Figure 1] FPTPGBs and its improved structure have been extensively investigated because of their particular advantages over traditional tilting pad gas bearings (5, 13-15, 17-20). San Andr\u00e9s (13) built a computational model, which considering the external pressurisation, to predict the static and dynamic performance of hybrid FPTPGBs. The model was confirmed by the experimental results. Zhu and San Andr\u00e9s (14) investigated the rotordynamic performance of pressurized FPTPGBs. Their experimental results showed that the rotordynamic performance of the bearings is superior to that of pressurized three-lobe cylindrical bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000409_aero.2015.7118995-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000409_aero.2015.7118995-Figure2-1.png", + "caption": "Figure 2. Schematic bearing with spall, explaining the unloading of the ball in the spall and subsequent impact at the spall trailing edge.", + "texts": [ + " [6] performed stress analysis based on those results, trying to understand the damage propagation mechanism. An example of damage propagation from the test performed by Rosado et al. [5], is shown in Figure 1. The results of the endurance experiments and the stress analysis indicated that spall propagation is a three step process: 1. The spall is broadened first axially across the width of the raceway, as shown in Fig. 1(b), until it reaches the entire width of the raceway, as shown in Fig. 1(c). 2. At some point, when the spall is wide enough, the ball descends into the spall as illustrated in Fig 2. This causes the other balls to share the load that was supported by the descended ball. 3. The descended ball is pushed along the spall by the cage and impact the spall trailing edge. The impact results in severe fatigue damage that leads to the propagation of the spall in the circumferential direction of the raceway (Fig. 1(d) and Fig. 2). The main goal of the research is to build prognostic tools for remaining useful life (RUL) of the bearings. It is essential, for estimation of RUL, to be able to model accurately the damage initiation and propagation in bearings. In the current study a model that simulates damage initiation is presented. The next sections present the contact boundary conditions, the damage initiation model, and its implementation results. The contact between the rolling element and the raceway is represented by Hertzian contact model [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.16-1.png", + "caption": "Figure 10.16 Leg details of the Stewart\u2013Gough platform.", + "texts": [ + "314) give \ud835\udf19k as follows without any additional sign variable. \ud835\udf19k = atan2(\ud835\udf0ekrk2, \ud835\udf0ekrk1) (10.315) According to the preceding equations, if \ud835\udf0ek = + 1 leads to \ud835\udf03k > 0 and \ud835\udf19k , then \ud835\udf0ek = \u2212 1 leads to \ud835\udf03\u2032k = \u2212\ud835\udf03k with \ud835\udf19\u2032 k = \ud835\udf19k \u00b1 \ud835\udf0b. However, the pairs {\ud835\udf03k , \ud835\udf19k} and {\ud835\udf03\u2032k , \ud835\udf19 \u2032 k} are not visually distinguishable. Therefore, the value of \ud835\udf0ek is not very significant. Nevertheless, \ud835\udf19k is preferred to be an acute angle so that c\ud835\udf19k > 0 in order to prevent the pedestal of the universal joint, which is shown in Figure 10.16, from unnecessarily excessive rotations. In other words, it is preferred to have |\ud835\udf19k|<\ud835\udf0b/2 instead of |\ud835\udf19k|<\ud835\udf0b. Equation (10.314) implies that this preference can be realized with \ud835\udf0ek = sgn(rk1). As for the angles of the spherical joint of the leg Lk , they can be found from Eq. (10.301) by rearranging it as shown below. eu\u03033\ud835\udf19 \u2032 k eu\u03032\ud835\udf03 \u2032 k eu\u03033\ud835\udf13 \u2032 k = C\u0302\u2217 k (10.316) In Eq. (10.316), the matrix C\u0302\u2217 k is known with the following expression. C\u0302\u2217 k = e\u2212u\u03032\ud835\udf03k e\u2212u\u03033(\ud835\udf19k+\ud835\udefdk )C\u0302eu\u03033\ud835\udefek (10.317) Equation (10", + " for known matrices v, \ud835\udf14, and C\u0302, the corresponding rates of the active and passive joint variables can be found leg by leg as explained below. For the leg Lk , Eq. (10.391) can be written as follows: u1(sk ?\u0307?k) + u2(sk?\u0307?ks\ud835\udf03k) + u3s\u0307k = wk (10.392) In Eq. (10.392), the column matrix wk is known with the following expression. wk = e\u2212u\u03032\ud835\udf03k e\u2212u\u03033(\ud835\udf19k+\ud835\udefdk )[v \u2212 ?\u0303?C\u0302(h0u3 \u2212 d0eu\u03033\ud835\udefek u1)] (10.393) Note that wk represents the velocity of the joint center Ak in the leg frame k(Bk) attached to Lk . This frame is formed by the basis vectors u\u20d7(k) 1 , u\u20d7(k) 2 , and u\u20d7(k) 3 as illustrated in Figure 10.16. Note also that sk > 0 always. Therefore, if s\ud835\udf03k \u2260 0, Eq. (10.392) leads to the following solution, in which wki = ut i wk for i = 1, 2, 3. ?\u0307?k = wk1\u2215sk (10.394) ?\u0307?k = wk2\u2215(sks\ud835\udf03k) (10.395) s\u0307k = wk3 (10.396) As for the rates of the angles of the spherical joint of Lk , they can be found from Eq. (10.390) after rearranging it as shown below. ?\u0307?\u2032 ku3 + ?\u0307?\u2032ku2 + ?\u0307? \u2032 keu\u03032\ud835\udf03 \u2032 k u3 = \ud835\udf14 \u2217 k \u21d2 u1(?\u0307? \u2032 ks\ud835\udf03\u2032k) + u2?\u0307? \u2032 k + u3(?\u0307?\u2032 k + ?\u0307? \u2032 kc\ud835\udf03\u2032k) = \ud835\udf14 \u2217 k (10.397) In Eq. (10.397), \ud835\udf14\u2217 k is known with the following expression" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002817_j.apm.2020.01.007-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002817_j.apm.2020.01.007-Figure6-1.png", + "caption": "Fig. 6. Characteristic curves in the coupler link.", + "texts": [ + "2502 \u00b0, \u03b8P 1 = 223.5878 \u00b0) and ( \u03b4P 2 = 82.2073 \u00b0, \u03b8P 2 = 348.6 6 69 \u00b0). Through the coordinate transformation Eq. (30) , coordinates of the two points in { R O ; i m , j m , k m } are ( x Bm1 = \u22120.2921, y Bm1 = \u22120.1508, z Bm1 = 0.94 4 42) and ( x Bm2 = 0.2682, y Bm2 = \u22120.9173, z Bm2 = \u22120.2942). Based on the modelling procedure and the programming, the spherical moving centrode, geodesic inflection curve and the spherical Ball curve are drawn in the link BC of the spherical linkage, as shown in Fig. 6 . It is worth mentioning that the whole spherical Ball curve is drawn including the symmetric part. It is obvious that the moving centrode and spherical Ball curve are both envelopes of the family of geodesic inflection curves. For a spherical four-bar linkage, the shapes of the coupler curves are various depending on the local and global characteristics of the curves. The local properties are reflected by the geodesic curvature of the coupler curve, while the global characteristics are difficult to describe for lack of similar kinematic invariants" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003929_s40430-020-02785-6-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003929_s40430-020-02785-6-Figure1-1.png", + "caption": "Fig. 1 Three structural schemes of WSB", + "texts": [ + " And the ISFD technology is introduced into the conventional water-lubricated stern bearing (CWSB) to achieve a waterlubricated damper stern bearing (WDSB). The finite element method (FEM) is used to calculate the harmonic response of these two bearings, and the damping structure parameters are optimized. The dynamic characteristics of two bearings are measured on a bearing test rig to verify the vibration reduction effect of WDSB. A CWSB consists of a liner and a bushing, where the lining is made of polymer composite materials such as nitrile rubber (NBR). Figure\u00a01 presents three kinds of structure schemes. Among them, the design using rubber to carry the load and reduce vibration at the same time remains defective. The dynamic model of CWSB is shown in Fig.\u00a02a, where m is the vibration mass, k and c are the stiffness and damping coefficients of the bearing, respectively, which are formed by the lining and the water film. Under the excitation force F, the shaft will give response x. Under the heavy load, an obvious deformation of rubber layer in lining will come up, and the damping effect of rubber will be significantly reduced after a serious extrusion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003194_s12206-020-0503-y-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003194_s12206-020-0503-y-Figure2-1.png", + "caption": "Fig. 2. Structure of the new-type active TDB.", + "texts": [ + " The dynamic responses after rotor drops onto the new-type TDB are theoretically and experimentally analyzed. The structure of the AMB system equipped with the newtype active TDB in this study is shown in Fig. 1. The new-type TDB is assembled in the left end, while the right end still uses traditional TDB. The magnetic bearing controller (MBC) controls each magnetic bearing to levitate the rotor in the center of the stator. The rotor can conduct noncontact rotation around its axial direction. The structure of the new-type TDB in this study is shown in Fig. 2. The ball bearing is mounted on the rotor and rotates with the rotor. Two fixed supporting bases (FSB) are arranged in the vertical direction with the fixed protective clearance. Two moveable supporting bases (MSB) are arranged in the horizontal direction. During normal operation of AMB, the MSB keeps the protective clearance unchanged under the elastic force from the support spring. In case of the AMB failure, the MSB will eliminate the protective clearance under the action of the electromagnetic force from the two electromagnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001178_ssd.2013.6564123-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001178_ssd.2013.6564123-Figure1-1.png", + "caption": "Fig. 1. Inverted pendulum system", + "texts": [ + " Now, the second subsystem will be considered which can be represented by this following form: 2 2 2SS S LESx f x g x u y h x where 2 1 T S m nf f f f and 1 T m ng g g In the application of the technique of input-output feedback linearization, the control law LESu is given by the following expression: 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 60 61 2 2 2 1 S S S r f LES r g f v L h x u L L h x Finally, the control law is designed: 2 2 2 1 1 1 1 S S S r f m m m m m m r m g f v L h x u x f g e k e g L L h x The proposed algorithm is composed of the following 4 steps: Step 1: The first stage of the control strategy which we proposed is the Decomposition of the system (19) in two subsystems 1S and 2S respectively of order 1n m and 2n n m Step 2: Synthesis of a control law for subsystem 1S represented by the equation (21) uBac using the backstepping technique. Step 3: Synthesis of a control law for subsystem 2S represented by the equation (22) LESu using the exact inputoutput linearization. Step 4: Finally the control law overall system is synthesized by the sum of two control laws represented by (34) and (36) IV.APPLICATION TO CONTROL OF NON LINEAR MODEL OF INVERTED PENDULUM A. Mathematical Model of Inverted Pendulum We study, in this section, a nonlinear model of an inverted pendulum [16] (fig.1). This is an ideal model to verify a new control theory or methods as an absolute instability, high, variable and high coupling typical nonlinear system [17]. The design objective is to stabilize the pendulum in its unstable equilibrium point. From the Lagragien, the equations of motion are: 2cos sin cos sin 0 p p M m y ml ml u l y g ml (38) where is the angle of the pendulum, py is the displacement of the cart, and u is the control force, parallel to the rail, applied to the cart" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000014_iros40897.2019.8967565-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000014_iros40897.2019.8967565-Figure3-1.png", + "caption": "Figure 3. Rolling without slipping for the elliptic foot.", + "texts": [ + " An ellipse in this plane is made up of points of the form: (4) where ra and rb are the ellipse\u2019s major and minor radii on the xE and yE axes respectively, and \u03c6c is a parameter visualized as the angle of a line whose intersections with two circles of radius ra and rb result in the (xi, yi) coordinates of each ellipse point, as is shown in Fig. 2. This point on the ellipse determines the geometric angle \u03c6e, for which: (5) This last equation links a geometric point (xi, yi) of the ellipse\u2019s perimeter to the parameter \u03c6c used in the definition of the ellipse, enabling the expression of contact geometry through simple algebraic equations. C. Contact geometry in terms of state variables In Fig. 3, the elliptic foot is tangent to the ground at the point of contact. Therefore, for a foot rotation angle \u03b8, see Fig. 3, the tangent to the ellipse at the contact point, measured in the xE-yE CS, can be calculated using the definition of the tangent line for Eq. (4): (6) providing a relationship between \u03c6c and \u03b8. Therefore, at the contact point, \u03c6e can be linked to the leg angle \u03b8 using (5): (7) D. The kinematics of rolling on elliptic feet The elliptic foot of Fig. 3 performs rolling without slipping on flat ground: by increasing the leg angle from 0 to q = \u03b8 , L1,\u03c8 , L2\u23a1\u23a3 \u23a4\u23a6 T !q x = q; !q\u23a1\u23a3 \u23a4\u23a6 = \u03b8 , L1,\u03c8 , L2 , !\u03b8 , !L1, !\u03c8 , !L2\u23a1\u23a3 \u23a4\u23a6 T xr = r\u03b8 E (xi , yi ) = (ra cos\u03d5c ,rb sin\u03d5c ) tan\u03d5e = yi xi = rb sin\u03d5c ra cos\u03d5c = rb ra tan\u03d5c ( cos ) tan ( sin ) tan E i i c b c b i c i a c a c dy dy d r r dx d dx r r j j q j j j - = = = = - tan\u03b8 = \u2212rb 2 ra 2 tan\u03d5e 6303 Authorized licensed use limited to: Western Sydney University. Downloaded on July 26,2020 at 08:24:08 UTC from IEEE Xplore. Restrictions apply. \u03b8, the resulting displacement of the contact point, xroll must be equal to the ellipse\u2019s arc length xe: Knowing xroll, it is possible to express the position of the center of the ellipse, C, in x-y coordinates: (9) (10) In (9), x0 is the distance of the contact point from the origin when \u03b8=0, as is shown in Fig. 3. Note that \u03c6e, \u03c6c<0 for the bottom part of the ellipse, where contact occurs. The spatial configuration of the rest of the biped can be expressed with respect to point C\u2019s coordinates. Note that the x-y CS is rotated by a<0 with respect to the global CS; therefore, both xc and yc are used for the estimation of the biped\u2019s potential energy due to gravity g in the X-Y CS. The Lagrangian will then contain the integral xroll, introduced through the potential energy terms. This integral does not have an analytical solution; however, only the time and state derivatives of xroll are used to produce the dynamic equations of the biped in the Lagrangian formulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000191_b978-0-08-100543-9.00018-x-Figure18.15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000191_b978-0-08-100543-9.00018-x-Figure18.15-1.png", + "caption": "FIG. 18.15 VAMI electrolyzer with a bulk anode for producing titanium powder.", + "texts": [ + " Raw materials for anodes include substandard titanium sponge, alloys based on titanium, and conducting products of titanium raw material reduction (oxycarbides, oxycarbonitrides, etc.). The electrolyte is the molten NaCl, or NaCl + KC1, or NaCI + KC1 +MgCl2. The preparation of the electrolyte consists in the tetrachloride \u0422iCl4 reduction in molten NaCI and KCI using titanium scrap or sodium. Titanium ions are reduced to metal stepwise on the cathode; on the anode titanium is oxidized up to Ti2+ and Ti3+. An average degree of oxidation (valency of titanium ions) in the electrolyte is 2.2\u20132.3. The process of titanium refining was carried out in electrolyzers of VAMI design (Fig. 18.15). The electrolyzer is a two-cell device with a common heating furnace. Each cell consists of a retort with bulk anode containers and a cathode chamber hermetically connected with the retort. Both cells are electrically connected in parallel. The cathode chamber is supplied with mechanisms for moving the cathode and for cutting off the deposit as well as the device for additional loading of the anode material, which consists of bunkers with feeders and discharge spouts with shutters. The device for unloading the anode material in the form of a mobile tray with a folding bottom and the metal receiver is mounted on the retort" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003969_0954406220983367-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003969_0954406220983367-Figure15-1.png", + "caption": "Figure 15. Finite element model (a) boundary conditions of rough surface; (b) local mesh refinement of rough surface.", + "texts": [ + " A Real contact area after deformation m2\u00f0 \u00de A0 Real contact truncation area before deformation m2\u00f0 \u00de A1 Contact area of slider and roller m2\u00f0 \u00de A2 Contact area of guideway and roller m2\u00f0 \u00de Al Maximum contact area of asperity m2\u00f0 \u00de Apc Critical contact area of plastic deformation m2\u00f0 \u00de Aec Critical contact area of elastic deformation m2\u00f0 \u00de An Nominal contact area m2\u00f0 \u00de D fractal dimension E comprehensive elastic modulus MPa\u00f0 \u00de F friction factor f1 n\u00f0 \u00de middle function f2 n\u00f0 \u00de middle function Fc total contact load N\u00f0 \u00de Fe contact load under elastic deformation N\u00f0 \u00de Fep contact load of elastic-plastic deformation N\u00f0 \u00de Fp contact load under plastic deformation N\u00f0 \u00de g1 D\u00f0 \u00de constant associated with the fractal dimension g2; g3; g4 constants associated with the fractal dimension G fractal roughness K D\u00f0 \u00de combination parameter Kf friction correction factor l transverse width of section before asperity deformation m\u00f0 \u00de l0 Transverse width of section after asperity deformation m\u00f0 \u00de le Effective contact length of roller m\u00f0 \u00de L Sampling length m\u00f0 \u00de M Number of superimposed ridges n Integer from 1 to 4 n1 Low cutoff frequency n\u00f0A\u00de Number distribution function of contact asperities pe Contact pressure under elastic deformation MPa\u00f0 \u00de Pep Contact load under elastic-plastic deformation N\u00f0 \u00de pmax Maximum contact pressure of asperity MPa\u00f0 \u00de Pp Contact pressure of plastic deformation MPa\u00f0 \u00de P Dimensionless contact load N\u00f0 \u00de Q Spatial frequency index R Curve radius at the top of asperity um\u00f0 \u00de Ra Arithmetic average height um\u00f0 \u00de Rq Root-mean-square (rms) roughness um\u00f0 \u00de R Equivalent curvature radius um\u00f0 \u00de S w\u00f0 \u00de Power spectrum w Linear load N\u00f0 \u00de Xn Index z x; y\u00f0 \u00de Three-dimension rough surface profile um\u00f0 \u00de zw x; y\u00f0 \u00de Two-dimension rough surface profile um\u00f0 \u00de * Dimensionless symbol C Frequency density of rough surface profile d Deformation amount at the top of asperity um\u00f0 \u00de dec Critical elastic deformation of asperity m\u00f0 \u00de ry Yield strength MPa\u00f0 \u00de q radius of curvature of the asperityX A Sum contact area of two solid surfaces m2\u00f0 \u00de s Solid surface contact coefficient t Poisson\u2019s ratio / Material characteristic coefficient /m;q Phase Through got fractal parameters in above section 3.2, the geometric micro-topography can be determined. Generally, the geometrically complex contacting problem has to be discretized by adopting the finite element method, such as the asperities model of rough surfaces in this paper. These asperities peaks and valleys are expected to deform elastically and plastically under loading, therefore the ABAQUSVR model (finite element model) contains three key points in present contact analysis as follow. Figure 15 shows the mesh discretization used in the model consisting of the two contacting bodies, namely rough surface and rigid plane. The rough sur- face is discretized using triangular and quad elements; moreover the current model is divided into four layers of different mesh sizes by using transition elements. That will be more efficient in iteration calculation. The local asperity contact model is shown in Figure 15(b). Special attention is paid to the mesh of the asperities, with a total number of 25,341 elements in the rough surface. Triangular elements are chosen as the transition elements and quadrilateral elements similar to the other regions. The number of elements varies with the rough profile size changes. For the rough profile model in Figure 15(a), a mesh of 24,979 linear quadrilateral elements of CPS4R type and 362 linear triangular elements of type CPS3 type is generated. The material of slider and guideway is GCr15 widely used in the manufacturing industry, which belongs to a sort of high carbon chromium bearing steel with less alloy content. Therefore, the mechanical performance of bearing steel GCr15 is pivotal to the research of elastic and plastic deformation under loading. After quenching and low-temperature tempering, the properties of high hardness, uniform structure, good wear resistance and high contact fatigue will be obtained. As shown in Table 1, the elastic modulus is 2:1 105MPa, Poisson\u2019s ratio is 0.3, hardness is 58- 62HRC. From Figure 16, the yield strength of bearing steel GCr15 is approximately 518MPa, the true stress-strain curve of the material GCr15 is as follows:57 From the previous Figure 15, the plastic deformation of material GCr15 probably occurs in the contact asperities peaks, where asperities reach the yield strength limit if given enough load. Hence, plastic strain calculation can be expressed as epl \u00bc et eel \u00bc et r=E (45) where eplis true plastic strain, etis total true strain, eelis true elastic strain. The contact feature of rough surface is actually the contact of rough peaks in the microscopic scale. The contact calculation of rough peaks includes geometric nonlinearity, contact nonlinearity and material nonlinearity", + " So the finite element analysis of rough surface is based on nonlinear contact analysis. For the contact model between rough surface and the rigid plane, the condition is achieved by fixing the rigid plane completely, and the displacement for the top nodes and both lateral nodes of the rough surface is restricted in x direction. Meanwhile the normal pressure value 4:5 10 4MPa is enforced on the top nodes of the rough surface, and the interaction is built between the contact surfaces, which mean that all nodes of the rough surface are free in y direction, as shown in Figure 15(a). Results of finite element analysis between rough surfaces The contact stress and strain results of the rough surface compressed by the rigid plane are shown in Figure 17(a) and (b). All peaks of asperities are compressed into severe deformation, however, other valleys are less deformed. The maximum stress value is 990MPa shown in Figure 17(a), which exceeds the yield strength limit of material GCr15, the maximum displacement value is 0.3mm at the normal pressure 4:5 10 4MPa shown in Figure 17(b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002946_j.precisioneng.2020.03.001-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002946_j.precisioneng.2020.03.001-Figure16-1.png", + "caption": "Fig. 16. Setup of the experiment for measuring contact stiffness: (a) Horizontal contact stiffness, (b) Vertical contact stiffness.", + "texts": [ + " An experiment was conducted to measure changes in the contact stiffness. In this experiment, contact surfaces with different distributions of the real contact areas were made using the CMC method. The contact stiffness of each contact surface was measured for comparison. The specimens were pushed onto a base block to make a contact surface. Fig. 15 shows a schematic of the top specimens and the base block. The specimens had different distributions of the real contact areas because the specimens had different amplitudes of waviness. Fig. 16 shows the experimental setup for measuring contact stiffness. The horizontal and vertical contact stiffness was measured for each specimen. A vertical preload was applied using a bolt and measured using force sensor 1. Horizontal and vertical forces were applied to the specimen using a piezoelectric element. The vertical force was measured with a force sensor 1 and the vertical force was measured with the force sensor 2. The relative displacement between the top specimen and the base block was measured with a capacitance-type displacement sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001303_0954406213515644-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001303_0954406213515644-Figure5-1.png", + "caption": "Figure 5. Simplified micro-contact model for a single asperity.", + "texts": [ + "comDownloaded from Micro-contact model of metal surfaces Figure 3 shows the metal contact-sealing structure of the globe valve. From a micro-scale view, metal seal contact can be reduced to the contact between two elastic rough surfaces, and the micro-contact model is shown in Figure 4. As a consequence, the microcontact between the two flat surfaces, in fact, is regarded as the contact between asperities of rough surfaces. For a single asperity contact, the interfacial profile can be approximately taken as a spherical surface. The micro-contact state between two asperities is shown in Figure 5, where R1 and R2 are the interfacial curvature radii of two asperities, respectively, i is the mutual approach, Wi the external load and Ai the contact area for a couple of contacting asperities. The curvature radius R of a single asperity is calculated as follows: the contact surface is divided into uniform grids in the directions X and Y (shown in Figure 6); the value of height profile for any interfacial point is calculated from the modelled surface of Figure 1. In order to calculate the curvature radius Ri,j of any interfacial point (i,j) in rough surface, we can calculate the curvature radii of the point (i,j) in the directions X and Y, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000945_acc.2013.6580710-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000945_acc.2013.6580710-Figure1-1.png", + "caption": "Fig. 1 : Blending Schemes of a Dual Missile", + "texts": [], + "surrounding_texts": [ + "Washington, DC, USA, June 17-19, 2013 Force and Moment Blending Control for Agile Dual Missiles Seunghyun Kim, Dongsoo Cho and H. Jin Kim Ahstract- This paper presents a force and moment blending control scheme for a dual-controlled missile with tail fins and reaction jets, which is especially aimed at a fast response. A dual missile can be controlled by forces and moments independently, when the reaction jet is located in front of the center of gravity of the missile. Controlling the net force of aerodynamic lift and jet thrust yields much a faster response compared with controlling the net moment, but it uses large control efforts caused by divergence of angle of attack, especially the jet thrust. Here, force control is implemented by using sliding mode control. Large input usage is alleviated by shaping angle of attack. The proposed blending scheme begins with force control and then makes a transition to moment control. Both control strategies are demonstrated by a nonlinear missile system. The smooth transition from force control to moment control is demonstrated. The proposed approach shows a very fast response, while its input usage is almost same as the conventional moment control." + ] + }, + { + "image_filename": "designv11_34_0000207_9781119454816.ch8-Figure8.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000207_9781119454816.ch8-Figure8.4-1.png", + "caption": "Figure 8.4 Four-bar linkage.", + "texts": [ + " (E8) to (E11) gives \u03bb\u221701 = 1, \u03bb\u221711 = 1, \u03bb\u221721 = 1 2 , and \u03bb\u221722 = 1 2 . Thus, the maximum value of v and the minimum value of f is given by v\u2217 = (0.188 1 )1 (1.75)1 (900 0.5 )0.5( 1 0.5 )0.5 (0.5 + 0.5)0.5+0.5 = 19.8 = f \u2217 The optimum values of xi can be found from Eqs. (8.62) and (8.63): 1 = 0.188y\u2217d\u2217 19.8 1 = 1.75y\u2217h\u2217\u22121d\u2217\u22121 1 2 = 900y\u2217\u22122 1 2 = y\u2217\u22122h\u22172 These equations give the solution: y* = 42.426, h* = 30 in., and d* = 2.475 in. Example 8.14 Design of a Four-Bar Mechanism Find the link lengths of the four-bar linkage shown in Figure 8.4 for minimum structural error [8.24]. SOLUTION Let a, b, c, and d denote the link lengths, \ud835\udf03 the input angle, and\ud835\udf19 the output angle of the mechanism. The loop closure equation of the linkage can be expressed as 2ad cos \ud835\udf03 \u2212 2cd cos\ud835\udf19 + (a2 \u2212 b2 + c2 + d2) \u2212 2ac cos(\ud835\udf03 \u2212 \ud835\udf19) = 0 (E1) In function-generating linkages, the value of \ud835\udf19 generated by the mechanism is made equal to the desired value, \ud835\udf19d, only at some values of \ud835\udf03. These are known as precision points. In general, for arbitrary values of the link lengths, the actual output angle (\ud835\udf19i) generated for a particular input angle (\ud835\udf03i) involves some error (\ud835\udf00i) compared to the desired value (\ud835\udf19di), so that \ud835\udf19i = \ud835\udf19di + \ud835\udf00i (E2) where \ud835\udf00i is called the structural error at \ud835\udf03i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003985_s10514-020-09962-5-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003985_s10514-020-09962-5-Figure1-1.png", + "caption": "Fig. 1 Leader\u2013follower scheme", + "texts": [ + " Note that if the lateral velocity is cancelled, i.e., vyp = 0, then, it is possible to obtain the kinematic model of the differential-drive mobile robot, given by x\u0307 = vx cos \u03b8 , (2a) y\u0307 = vx sin \u03b8 , (2b) \u03b8\u0307 = w , (2c) for all \u2208 ND. It is worth mentioning that system (2) is underactuated and satisfies the following non-holonomic restriction x\u0307 sin \u03b8 \u2212 y\u0307 cos \u03b8 = 0. Let us recall the dynamic model, based on distance and orientation, between two mobile robots (Gonz\u00e1lez-Sierra et al. 2018; Paniagua-Contro et al. 2019), defined as (Fig. 1) \u03b7\u0307 j i = f\u03b7 j i (\u03b7 j i )u j + g\u03b7 j i (\u03b7 j i )ui , i = j, i, j \u2208 N , (3) where \u03b7 j i = [ d ji \u03b1 j i \u03b8i ] \u2208 R 3 is the state vector, d ji \u2208 R+ is the distancemeasured from the geometrical center of agent R j to the geometrical center of the agent Ri , with R+ as the set of all positive real numbers, d ji x and d ji y \u2208 R+ are the components of the distance vectord j i with respect to a global frame and \u03b1 j i \u2208 R is the formation angle measured from the distance vector d j i to a local frame attached to the agent Ri " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001315_j.mechmachtheory.2015.02.002-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001315_j.mechmachtheory.2015.02.002-Figure7-1.png", + "caption": "Fig. 7. The equivalent open-loop chain for the GRM.", + "texts": [ + " Absolute angular velocities of the GRM links In order to obtain the absolute angular velocities of the GRM links, the graph shown in Fig. 5 is used. From Fig. 5 the following absolute angular velocity equations can be written: \u03c90 1 \u00bc \u03c90 10 \u03c90 4 \u00bc \u03c90 40 \u03c90 5 \u00bc \u03c90 50 \u03c90 6 \u00bc \u03c90 61 \u00fe\u03c90 10 \u03c90 2 \u00bc \u03c90 21 \u00fe\u03c90 10 \u03c90 3 \u00bc \u03c90 32 \u00fe\u03c90 21 \u00fe\u03c90 10 \u03c90 7 \u00bc \u03c90 72 \u00fe\u03c90 21 \u00fe\u03c90 10: \u00f047\u00de Denavit\u2013Hartenberg (D\u2013H) convention [4] is used in order to obtain \u03c921 0 , \u03c932 0 , \u03c961 0 and \u03c972 0 relative to the base frame. First, a coordinate frame is assigned to each link of the equivalent open-loop chain of the GRM based on the D\u2013H convention as depicted in Fig. 7. Then, link parameters (D\u2013H parameters) are determined between the successive frames. Actually in D\u2013H convention, the aim is to find the end effector position and orientation relative to base coordinate frame. For the GRM shown in Fig. 7, (x3y3z3), (x2y2z2), and (x1y1z1) frames attached to links 3, 2, and 1, respectively, are sufficient to determine the position and orientation of the end effector. However, since the absolute angular velocities of the other links are determined by using the Matroid method, two additional coordinate frames are required in order to determine the other links' absolute angular velocities as well. Link 4 is rotating about the z1\u2032 which is parallel to the z0, link 5 is rotating about z0, link 6 is rotating about z1 and link 7 is rotating about z2\u2032 The link parameters of the GRM are shown in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002558_0954410016680644-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002558_0954410016680644-Figure1-1.png", + "caption": "Figure 1. Position coordination variable for three UAVs: (a) unsteady case; (b) steady case.", + "texts": [ + " Suppose that the predefined geometric configuration is known, and then the predefined position vector from Ui to a reference point for the formation is determined, which is invariable and defined as qd iF. Therefore, each UAV generates a reference point qiF, and qiF is chosen as the position coordination variable. Alternatively, the position coordination variable can be obtained by vector computation: qiF \u00bc qi \u00fe qd iF \u00f01\u00de The position coordination variables for three UAVs are demonstrated as in Figure 1 in both the unsteady case and the steady case, where the small circles represent the position coordination variables. For brevity and clarity, only the position coordination variable q2F is clearly marked. If qiF! qiF as t!1, the formation keeping is realized as in Figure 1(b). Multi-UAV formation control strategy A cooperative guidance system and a cooperative control system, which aim to control the position at University of Warwick on December 11, 2016pig.sagepub.comDownloaded from and attitude, respectively, constitute the multiUAV formation control strategy. The cooperative control system is constructed as the inner loop, and the cooperative guidance system is constructed as the outer loop, where the output of the outer loop is used as the input of the inner loop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000338_1077546315586495-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000338_1077546315586495-Figure1-1.png", + "caption": "Figure 1. Dynamic model of the half rear suspension-tire-tread system.", + "texts": [ + " Then, through bifurcation analysis of the tire tread we found that the speed of the vehicle and slip angle of wheel play significant roles in vibration generation. For investigating the influence of this self-excited vibration on the system, the system model has been simulated with different parameters, such as vehicle speed, vertical load, and tire pressure. The result explains other phenomenon characteristics mentioned above. In view of the features of polygonal wear on tires, the rear suspension-driven-tire system has been simplified as shown in Figure 1. The assumptions are applied as follows. 1. It is a dynamic model of the half rear suspensiontire-tread system. Four Cartesian coordinate systems have been assumed to express the model. They are the coordinate system of the ground x y z\u00f0 \u00de 0, which represents the forward direction, lateral direction, and vertical direction; the coordinate system of the car body n01n02n03\u00f0 \u00de 0, whose directions are the same as the coordinate system of the ground, only its origin O keeps constant speed in the x direction; the coordinate system of the wheel n11n12n13\u00f0 \u00de 0, where the center of wheel G is also the origin; and the coordinate system of tire tread n21n22n23\u00f0 \u00de 0, taking the center of mass E of the tread contact block as the origin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003308_s40430-020-02488-y-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003308_s40430-020-02488-y-Figure10-1.png", + "caption": "Fig. 10 Gear meshing contact zone", + "texts": [ + " He [42] studied the solution method of the logarithmic spiral bevel gear contact zone. On the basis of the LSBG, the tooth surface was divided into the mesh and the coordinate values were obtained later. Then, the elliptic function method was combined to find the solution method of the contact area of the LSBG, and the boundary of the contact area was judged and cut. Finally, the contact area of the tooth surface of the LSBG was obtained in the finite element software, including position, size and shape, as shown in Fig.\u00a010. By this method, the cost of testing with a solid sample can be reduced. Wu [24] carried out meshing simulation analysis about LSBGs and studied the calculation methods of gear strength, including: tooth surface contact stress and root bending stress. The main innovation point was as follows: The previous tooth root stress calculation standards were based on the cantilever beam formula. The stress data of gear meshing were obtained by ANSYS simulation, after that the MATLAB curve fitting theory was used to fit the contact stress equation of the tooth surface and the bending stress equation Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:400 1 3 400 Page 8 of 14 of the tooth root by the least square method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003437_s00170-020-05886-7-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003437_s00170-020-05886-7-Figure11-1.png", + "caption": "Fig. 11 A concept of pressure-fed flexure mechanism [36]", + "texts": [ + " Positive hybrid effects exist by substituting carbon fibers with glass fibers, and applying a thin layer of GFRP on the compressive surface yields the highest flexural strength. Flexures are primarily constraint elements that utilize material elasticity to produce small, yet frictionless and precise inmotions, and an ideal flexure provides infinite stiffness and zero displacements along its degrees of constraint (DOC). Dynamic and thermal characteristics of the flexures are influenced by constitutive properties of materials, geometric compatibility, or force equilibrium conditions as illustrated in Fig. 11. In other words, AM technology that makes use of the flexibility in mechanical design can effectively control the material distribution (channel size, shape, and placement) in terms of stiffness, damping, resonance frequency, and heat dissipation of the flexures. In this study, the pressure-fed mechanism applying different pressure across the internal channel was firstly characterized by dynamic and thermal testing. The pressure distribution and direction around each channel surface will be dependent on the channel shape, and the dynamic characteristics of flexures will rely upon the crosssectional geometry of the channel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003515_j.ijsolstr.2020.08.022-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003515_j.ijsolstr.2020.08.022-Figure1-1.png", + "caption": "Fig. 1. A cantilever tensegrity.", + "texts": [ + " (34) and (35), the iterative equation for the error correction can be written as Ddq \u00f0s\u00de Dea 0 \u00f0s\u00de \" # \u00bc dq \u00f0s\u00de ea 0 \u00f0s\u00de \" # dq \u00f0s 1\u00de ea 0 \u00f0s 1\u00de \" # \u00bc D\u00f0s 1\u00de \u00fe F \u00f0s 1\u00de Dw c \u00f0s 1\u00de 2 64 3 75 \u00f036\u00de where D\u00f0s 1\u00de \u00fe is the Moore-Penrose pseudoinverse of D\u00f0s 1\u00de which takes the following form D\u00f0s 1\u00de \u00bc Kq T \u00f0s 1\u00de Ha\u00f0 \u00de\u00f0s 1\u00de Uq c \u00f0s 1\u00de Ua c \u00f0s 1\u00de \" # \u00f037\u00de The correction will be repeated using Eq. (36) until both k F k2 and kDw c k2 are smaller than the given tolerances. A cantilever tensegrity consisting of four stacked quadriprism modules named A, B, C, and D from left to right is shown in Fig. 1. In the initial configuration x0, the bottom (left) and top (right) surfaces of each module are squares with side lengths of a = 4 ffiffiffi 2 p m. For module A or C, the top surface is rotated by an angle of h \u00bc p=8 counterclockwise around the x-axis with respect to the bottom surface (Fig. 2 (a)). Module B or D is rotated clockwise by the same angle (Fig. 2(b)). Modules B, C and D are rotated clockwise around the x-axis by p=4, 0, and p=4, respectively, relative to module A, as shown in Fig. 2(c). Two adjacent modules are assembled by replacing the cables in their shared surfaces with saddle cables. Diagonal cables are then added to the side surfaces of each module. In addition, the side elements are arranged in opposite rotation directions for two adjacent modules. There are 32 joints, 16 bars and 72 cables in total in the cantilever tensegrity. The joint numbers are also shown in Fig. 1, and each member is denoted by the numbers of its two end joints. Six DOFs, A1 - x;A1 - y;A1 - z;A2 - x;A2 - y;A3 - x, are constrained on the vertical surface to prevent the rigid-body motion of the tensegrity. The height of each module is H = 4 m, and the overlap depth of adjacent modules is h. To be qualified as a tensegrity, a specified relationship must be satisfied between h and h=H (Nishimura, 2000). For the initial configuration shown in Fig. 1, it can be determined that h=H \u00bc 0:29509 corresponding to h \u00bc p=8. The member axial stiffness mk of all the bars in the tensegrity is 1 107 N, and that is 1 105 N for all the cables. It can be determined that the tensegrity has a unique mode of self-stress state (Pellegrino and Calladine, 1986). The axial forces of some typical members at the initial configuration x0 are listed in Table 1. The tensegrity will be deployed to a target configuration xm (Fig. 3) in which some DOFs of joints A6, A8, B6, B8, C6, C8, D6 and D8, which are all on the top surfaces of the modules, are specified to reach the given positions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002690_fit.2016.066-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002690_fit.2016.066-Figure1-1.png", + "caption": "Fig. 1. Aerial platform \u2013 Taurus", + "texts": [], + "surrounding_texts": [ + "Among the above mentioned inputs doublet and 3211 multistep were selected for exciting the lateral eigen modes of the UAV, as shown in Fig. 3. There are two options for input generation either apply it manually or we can generate automatic inputs. Manual inputs using RC flight controller involves human error causing inaccuracy. So automatic inputs were programmed using python scripts to get accurate movement of actuators. During flight experiment, these inputs were applied from the Ground Control Station through telemetry uplink." + ] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.18-1.png", + "caption": "Figure 10.18 illustrates the side view of the leg Lk of the manipulator together with the relevant geometric parameters and the active joint variable \ud835\udf03k . The side views of the other two legs are of course similar due to the symmetric configuration of the manipulator.", + "texts": [ + "440), \ud835\udf07\u2032 = 6 is the insignificant mobility associated with the spinning freedom of the lower links of the legs between the spherical joints and\ud835\udf07*= 3 is the significant mobility of the moving platform that is controlled by the rotary actuators of the three revolute joints. The preceding mobility analysis shows that the delta robot is made up of a deficient manipulator in the three-dimensional working space. Indeed, as it will also be verified in the forthcoming kinematic analysis, the parallel links of the lower legs prevent the moving platform from making any rotational motion. In other words, the moving platform can only move translationally with respect to the base frame 0(O). Incidentally, Figure 10.18 also illustrates the most common scenario, in which the delta robot is used. In this scenario, the fixed platform is situated at a certain height above the horizontal plane of the base frame and the delta robot works downward on the objects it handles. Referring to Figures 10.17 and 10.18, the following equations can be written in order to express the orientation of the moving platform (7) together with the end-effector and the location of the tip point (P) with respect to the base frame 0(O) by going through each leg" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001068_s1068375513030022-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001068_s1068375513030022-Figure1-1.png", + "caption": "Fig. 1. Scheme of ESA experimental facility for the mech anized coating. 1\u2014ALIER 31; 2\u2014vibrogenerator; 3\u2014 electrode; 4\u2014specimen; 5\u2014sliding carriage; 6\u2014regu lated drive of sliding carriage; 7\u2014installation bed; 8\u2014reg ulated drive of transversal vibrator.", + "texts": [ + " The parameters of technological pulses of the generator of an ALIER 31 installation are listed in Table 1. The operation mode 5 of that installation was used in the present study. The frequency of the preset pulses was ~0.1 kHz. This was reached using a special regulator of frequency (the energy coefficient). In order to produce nanofiber structures under controllable conditions of electrospark plating and determine the optimal modes, an experimental facility was developed for a mechanized coating with a wide range of parameters (Fig. 1). Standard vibrogenerator 2 of ALIER 31 was fixed on a vertical milling head of a milling machine so that it could perform oscillating movements with the adjustable amplitude and fre quency in the direction perpendicular to the move ment of sliding carriage 5 on which sample 4 was fas tened with screws. The amplitude was regulated by means of a special drive cam mounted on a vertical shaft of the machine with a possibility of adjusting it in the range of 1 to 10 mm. This made it possible to produce a track of coatings of different widths on the specimen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002042_tmag.2016.2527059-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002042_tmag.2016.2527059-Figure4-1.png", + "caption": "Fig. 4: Pointer diagram showing the forces acting on a rotor in stable levitation with no estimation error (a) and non-zero estimation error (b).", + "texts": [ + " The functionality of the observer block O can be illustrated with the following example. Assume that the rotor is levitating, the rotor angle is known, ie. \u2206\u03b8 = \u03b8\u0302 \u2212 \u03b8 = 0, and an external radial force ~Fr,e is applied. The net force acting on the rotor is zero in stable levitation. Therefore, the bearing force ~FB is pointing in the opposite direction of ~Fr,e. Inserting \u2206\u03b8 = 0 into (8) shows that ~F \u2217 B and ~IB are parallel to ~FB, note that kB is the bearing force constant. The corresponding pointer diagram is shown in Fig. 4(a). Figure 4(b) shows the same example with an angle estimation error \u2206\u03b8 6= 0. The net force acting upon the rotor is still zero, therefore ~FB is opposing ~Fr,e. However, F \u2217 B and subsequently ~IB are rotated by \u2206\u03b8 according to (8). The angle observer detects that the angle between ~Fr,e and ~IB is no longer 180 \u25e6 and corrects the angle estimate \u03b8\u0302 to drive \u2206\u03b8 to zero. Figure 5 shows a simulation of the operation of a bearingless machine operated with the angle estimator. The simulation models the machine shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002539_iecon.2013.6699805-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002539_iecon.2013.6699805-Figure2-1.png", + "caption": "Fig. 2. UPAT-TTR hardware setup", + "texts": [ + " Extendsive experimental studies verify the efficiency of the proposed controller and the improved agility on the longitudinal axis due to the full actuation that was gained by the rotor tilting mechanisms. The article is structured as follows: In Section II the UPATTTR platform is presented along with the extraction of a 978-1-4799-0224-8/13/$31.00 \u00a92013 IEEE 4174 system model. In Section III the attitude and the optimal translational controller are elaborated. In Section IV simulation results of the proposed control schemes and in Section V experimental results are presented. The article is concluded in Section VI. The UPAT-TTR platform setup is depicted in Figure 2. Its design is elaborated in [1]. The key characteristics enabling its autonomous operation are the following: a) A high computational power-capable Main Control Unit (MCU) based on the Intel Atom Z530 1.6 GHz CPU-based Kontron pITx Single Board Computer. b) An ARM Cortex M3-based microcontroller for the handling of sensor communication and control signals. c) A WiFi USB adapter for IEEE 802.11g based wireless communication. d) A Linux-based Operating System (OS), via which almost real-time execution control is achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002577_detc2016-59134-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002577_detc2016-59134-Figure7-1.png", + "caption": "Figure 7. Two types of atom fission. a) Fission type 1. The driver has +1DOF and the gripper has -1DOF. b) Fission type 2. The driver has +2DOF and the gripper has -2DOF.", + "texts": [ + "2 Stage 2: Atom fission In the second stage, the initial atom is fissioned into two counterparts \u2013 the driver and the gripper, with an equal value of positive and negative DOF, respectively. The 2D atom may be fissioned in two fashions. In fission type 1 the result is a pair of driver and gripper with +1 and -1 DOF, respectively. In fission type 2 the result is a pair of driver and gripper with +2 and -2 DOF, respectively. The driver (+DOF) is then withheld while the gripper (-DOF) is cast aside, to be used only in the fourth and final stage. Figure 7 shows the two types of 2D atom fission. 5.1.3 Stage 3: Addition of atoms In the third stage, new atoms are added onto the driver and onto each other. The first added atom has one connection to the driver, while its other connection is to the ground. From there on, every atom added must have at least one connection that is not grounded. There is no limit as to the amount of atoms which may be added in this stage. 5.1.4 Stage 4: The gripping Up until this stage, the graph had a positive DOF, that of the driver" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003656_s10015-020-00655-x-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003656_s10015-020-00655-x-Figure3-1.png", + "caption": "Fig. 3 Definition of the coordinate system of the tiltable coaxial rotor", + "texts": [ + " The definition of the rotation matrix WRB from the body coordinate system FB to the world coordinate system FW is described as below: in which RZ , RY and RX are the rotations around the axis ZB , YB and XB , respectively. The coordinate systems FP1 \u2236 { OP1 ;XP1 , YP1 , ZP1 } , FP2 \u2236 { OP2 ;XP2 , YP2 , ZP2 } , tilt angles 1 , 1 , 2 , 2 and some other variables of the coaxial rotors are also shown in Fig.\u00a02. For the expression of the model in a convenient way, the coordinate system of the i-th coaxial rotor ( i = 1, 2 ) is shown in Fig.\u00a03. The tiltable coaxial rotor-fixed coordinate systems lie in a plane and are separated by the angle of (see Fig.\u00a02), which are given by FPi \u2236 {OPi ;XPi , YPi , ZPi } (see Fig.\u00a03). Tilt angles (1)WRB = RZ( )RY ( )RX( ) i and i denote the ith coaxial rotor tilt angles about XPi and YPi , respectively. The initial tilt angles i and i are set to 0. The possible change about the tilt angles of the tiltable coaxial rotor is shown in Fig.\u00a04. In each tiltable coaxial rotor, the 360\u25e6 tilt mechanism is attached to the body coordinate system so as to change the pitch angle i of the coaxial rotor in 360\u25e6 (see Fig.\u00a04a), additionally, the 180\u25e6 tilt mechanism is connected to the coaxial rotors so that the roll angle i of the rotor can be changed in 180\u25e6 (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000116_077043-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000116_077043-Figure4-1.png", + "caption": "Figure 4. Macro photo injector.", + "texts": [], + "surrounding_texts": [ + "4th Annual Applied Science and Engineering Conference\nJournal of Physics: Conference Series 1402 (2019) 077043\nIOP Publishing\ndoi:10.1088/1742-6596/1402/7/077043\nThe supply of fuel flows through a high-pressure channel to the injector located on the cylinder head. The function of the injector is to atomize the fuel into the combustion system and how it works the injector has a valve that opens when the fuel pressure is high enough [2]. When the valve opens, the fuel is atomized and sprayed into the combustion chamber. At the end of the spraying, the pressure drops drastically and causes the valve to close again.\nThe electronic unit injection (EUI) system uses several of the same components as those used by the pump and line systems. The EUI system uses 1. Fuel tank, 2. Primary fuel filter, 3. Fuel transfer pump, 4. Secondary fuel filter, 5. Injector, 6. Return line, which is described in figure 3.\nA paper is high-pressure injector damage used on a common rail, an analysis is carried out of the cause of the damage and the results of a microscopic study of damaged components. Tribological damage from high-pressure injectors is local and cavitation holes. Cavitation position is mainly carried out on the valve, where the reduction in the amount of fuel injected [3].\nThe diesel engine used in trucks has problems when serviced. Inspection shows that four exhaust valves and intake valves and two exhaust valves and inlet valves are cracked. Fractographic studies show that fatigue is the main failure mechanism for all four valve springs. Under the action of the maximum nominal voltage, fatigue cracks begin in coil spring wire 1.3-1.5 from the top end of the", + "4th Annual Applied Science and Engineering Conference\nJournal of Physics: Conference Series 1402 (2019) 077043\nIOP Publishing\ndoi:10.1088/1742-6596/1402/7/077043\nspring. This region is also a location that is damaged by contact friction. The fracture of the intake valve and exhaust valve stem also indicate possible fatigue failure as a result of valve spring failure [4].\nMost recent diesel engine failures have occurred in modern diesel engines, which can be directly blamed on the quality of the fuel used. Due to poor fuel lubrication, as well as some particle contamination, the injector fails prematurely, causing bad combustion and engine damage, damage caused by fuel quality damage to the injector tip, abrasive on the plunger and barrel [5].\nThe deflection in the spring decreases with less stress when changing certain parameters. The spring index will affect deflection and stress. When the number of turns increases, the deflection decreases, and the pressure also decreases. This effect is not possible without changes in wire diameter because the wire diameter has an influence on deflection and pressure. Therefore when the wire diameter increases the deflection and the pressure also decreases. This happens because of the achievement of a better spring index [6].\nSpring is a mechanical shock absorbing system. Mechanical springs are defined as elastic objects that have the main function to bend or twist the load and to return to their original shape when the load is released. Researchers for many years have provided various research methods such as Theoretical, Numerical and Experimental. The researchers used theoretical, numerical and FEM methods. This study concludes that the Finite Element method is the best method for numerical solutions and calculates fatigue stress, life cycle and helical compression spring shear stress [7]. The helical compression spring is commonly used in diesel engine fuel injection systems, where it undergoes cyclic loading for more than 10 cycles. To predict the possible position of failure in the helical compression spring, which is used in the fuel injection system, along with the inner spring, finite element analysis is carried out, using ABAQUS 6.10. Simulation results show the behaviour of pressure oscillations along the length on the inside. It is ensured that oscillations are caused by bending associated with compression. It was also revealed that bending was caused by the geometry of a spring. The shear stress along the spring is found to be asymmetrical and with local maxims at the beginning of each middle coil. Asymmetry is caused by a coiled-coil smaller than 360 degrees.\nIn a visual inspection of a spiral spring injector, A \u00d8 wire diameter 7.0 mm and \u00d8 48 mm spring circumference has a broken first circle.", + "4th Annual Applied Science and Engineering Conference\nJournal of Physics: Conference Series 1402 (2019) 077043\nIOP Publishing\ndoi:10.1088/1742-6596/1402/7/077043\nPhoto macro image 8 on the injector spring the broken center point is a load bearing area and includes a fairly high center of stress concentration, it can be seen from the circumference area of the fracture there is a defect between the spring and spring wear.\nFigure 6 spiral spring circle 2 is the beginning of a fracture which first occurs the bending load thinning, while the shape of the fault has a 45o angle is a static fracture, while the 7 benchmark image on the injector spring fracture.\nMetallographic testing Figure 7 describes a benchmark that occurs in a spiral spring, that a benchmark indication is a form of fatigue in the injector spring." + ] + }, + { + "image_filename": "designv11_34_0001276_s12541-014-0425-7-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001276_s12541-014-0425-7-Figure2-1.png", + "caption": "Fig. 2 Body of RLW system", + "texts": [ + " RLW takes advantages of three main characteristics of laser welding: non-contact, single side joining technology, and high power beam capable of creating a joint in fractions of a second.12 These advantages have the potential to provide tremendous benefits on several fronts such as faster processing speed, reduced floor space, lower investment and operating costs, reduced tooling requirements, and significant reduction in energy usage and reduced environmental impact of vehicles.13 The RLW system is as shown in Fig. 2. In RLW systems, the laser beam is focused over the workpiece from a distance of about 0.5 m or more.14 A combination of mirrors and mechanical movement of the laser delivery mechanism results in very fast beam pointing. In fact, weld-to-weld repositioning may be less than 50 milliseconds. This is more efficient than traditional spot welding or more recent laser welding involving just robot motion because the seconds needed to move the robot from one weld to another are now eliminated.15 The quality of RLW output can be assessed visually which consists of three steps16 such as quality control before welding, quality control after welding, and quality reliability test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure2-1.png", + "caption": "Figure 2: Dimensions for test seals", + "texts": [ + "77 mm, which was equal to the rig rotor segment radius for the dynamic stiffness and leakage tests. The bristles were made of Haynes 25, which is typically preferred in most of turbine sealing applications due to their superior strength and satisfactory ductility up to 600oC. Cold worked Haynes 25 material with 10% cold reduction has 725 MPa tensile yield strength and 1070 MPa ultimate tensile strength limits at room temperature [10]. Seal photos are given in Figure 1, and brush seal dimensions are detailed in Figure 2. Using the axial spacing of 25% and tangential spacing value of 5%, the number of bristle rows in seal axial direction is found to be 15 by using the formula for a staggered bristle layout given by Aksit [7] (also visualized in Figure 3). 2 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use A free-state stiffness measurement system has been designed to characterize the seal stiffness under unpressurized static rotor conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002229_acc.2016.7525039-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002229_acc.2016.7525039-Figure6-1.png", + "caption": "Figure 6. Logarithmic norm of the complementary sensitivity function vs. logarithmic dimensionless frequency, \u03c4\u0304\u03c9, for different values of \u03b4 and \u03b5. \u03b3 = 0.3 and \u03bb \u03c4\u0304 = 0.4. The surfaces are plotted for three different levels of gain uncertainty, \u03b4 = 0,\u00b10.3.", + "texts": [ + " It is also obvious that the steady-state error requirement is met at low frequencies. To evaluate the performance of the closed-loop system with the proposed controller, let define M\u03b7 = \u2016\u03b7\u2016\u221e as the peak value of the complementary sensitivity function. As a rule of thumb, M\u03b7 is required to be less than 1.25 (0.1 dB) [31] or 1.05 (0.02 dB) [32]. For the present problem, this number is about 0.5 dB for the worst case scenario which happens for the case shown in Fig. 5 (see Fig. 3(a)). Similar large M\u03b7 observation applies for the other case within Fig. 6. Typically, a large value of M\u03b7 is an indication of a poor performance. In classical control, posing an upper bound on M\u03b7 has been a prevalent design specification. To improve the performance we will set M\u03b7 = 1.1 which gives the controller tuning parameter \u03bb\u03c4\u0304 = 1.5. The resultant complementary sensitivity function norms are illustrated through Figs. 7 and 8. The closed-loop response of the system with the proposed IMC method is presented in the next section. B. IMC Closed-Loop Time Response This section culminates in evaluation of the performance of the designed IMC by implementing it in a control loop subject to a unit step input" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003791_tase.2020.3035438-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003791_tase.2020.3035438-Figure1-1.png", + "caption": "Fig. 1. Compositions of the GSE system. The tested prostheses (HDP, prosthetic knee, and foot) are in the blue dotted oval.", + "texts": [ + " The objectives of this research are to solve the problems in the current HDPs research and tests, including: 1) lack of amputees for the prosthetic tests; 2) the safety of the hip disarticulation subjects could not be guaranteed at the preliminary research stage; and 3) quantitative and repeatable parameters could not be obtained systematically with the hip disarticulation subjects. Therefore, we developed a prosthetic gait simulation and evaluation (GSE) system that could simulate the pelvic track in natural human gait and drag the HDPs to walk on a treadmill. An adaptive neural network (ANN) control method was proposed to solve the problem of unstable motion and large trajectory tracking error when the manipulator was disturbed by the dragged HDPs. The GSE system, as shown in Fig. 1, contains a prosthetic thigh simulation robot (PTSR), a six-axis force sensor, and a pressure measuring treadmill. The treadmill was used to analyze the GRF. The PTSR for pelvic trajectory reproduction was six degrees of freedom (DOFs) manipulator (RT-Sim simulator, Beijing Links Co., Ltd.) and its D-H parameters were listed in Table I. In the structure of the system, as shown in Fig. 2, the trajectory data collected by the camera-based motion analyzer are smoothed by a Gaussian smoother. The smoothed trajectory is mapped into the PTSR workspace with an alignment of PTSR\u2019s initial position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003676_ecce44975.2020.9235378-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003676_ecce44975.2020.9235378-Figure11-1.png", + "caption": "Fig. 11. Constitutive elements of the test-bench.", + "texts": [ + " Using this test bench, a prototype of winding is first characterized and the possible deviations from the models are discussed and then a comparison of the performances of this prototype with a commercial actuator is provided. Comparing the performances of the two actuators requires a measure of the most relevant optimization criteria, namely the motor constant and the THD of the force. In this section, a description of the test bench is first provided. Then it is explained how the measure of the optimization criteria is achieved. Finally, the results are presented and discussed. The test bench that has been realized is shown in Fig. 10. It is made of several elements, labeled in Fig. 11 and each one of those elements are referenced in Table III Two identical rails of PM are symetrically placed on both sides of a linear guide on which a moveable cart is preconstrained. On this latter, both windings are fixed thanks to a support piece. This symmetric structure allows to place in competition the proposed winding and the commercial winding. For driving the motion, an IPOS8020 controller from 3619 Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 29,2021 at 10:13:24 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002692_mwscas.2016.7870019-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002692_mwscas.2016.7870019-Figure1-1.png", + "caption": "Fig. 1: Experimental magnetic levitation system.", + "texts": [], + "surrounding_texts": [ + "In this section, the efficacy of the proposed controller (12) is tested numerically on an experimental MAGLEV model shown in Fig. (1) [11] with specifications given in Table (I). MATLAB simulations are performed to verify the proposed controller performance and also, to compare it with the dynamic sliding mode controller (DSMC) [11]. The external perturbation in the system is considered as d = 0.5sin(0.2t). The proposed HOSTC parameters are k1 = 2, k2 = 1, k3 = 5, k4 = 2, \u03b2 = 0.5, \u03b3 = 1. The DSMC parameters are m1 = 15, m2 = 10, m3 = 7.5, \u03bb1 = 4.5, \u03bb2 = 4, \u03c4 = 8, w = 1.3. The simulation results obtained after application of both the controllers are presented in figures (2)-(5). Fig.(2) clearly verifies our claim of obtaining chattering-less and continuous control in case of the proposed HOSTC. Fig.(3) evidently shows that the proposed HOSTC outperforms the DSMC by providing fast convergence to the desired position (x1d = 0.01m) in finite time of 1.5sec which is almost 0.1 times faster than the DSMC\u2019s 15sec. It is more clear from fig.(3a) that the proposed HOSTC exhibit high accuracy and precision without any steady state error whereas fig. (3b) shows that DSMC is unable to stabilize the system at desired position. Fig.(4) depicts the velocity trajectory profile obtained from proposed HOSTC and DSMC. Fig.(4a)-(4b) clearly show that the proposed HOSTC helps the ball to stabilize faster at around 15sec than the DSMC is unable to stabilize even at 28sec. The coil current profile is shown in the Fig. (5). Figures (5a)-(5b) show that the current (1.1A) (peak) drawn by the coil in case of the proposed controller is very less than in the DSMC case (220A) (peak). Consequently, the size of the actuator will be reduced which will be beneficial in reducing the heating and saturation of the coil. Finally, in view of the simulation results obtained, it is verified that the proposed controller gives better performance result than the dynamic sliding mode control (DSMC) and hence can be applied to other practical systems. V. CONCLUSION The new high-order super twisting sliding mode controllers has been developed for the control of magnetic levitation system. It is observed that the proposed controller is effective in obtaining the high performance characteristics such as fast and finite-time stabilization, continuous and chattering free control, zero steady state error and better actuator performance. The proposed controller combines the attributes of homogeneity theory and conventional super twisting controller. The estimation of the state convergence time is obtained from the Lyapunov stabilty theorem. Simulation results have been included to prove the efficacy of the proposed control methods. REFERENCES [1] A. Goel and A. Swarup, \u201cHigh order super twisting sliding mode control of robotic manipulator,\u201d ICIC express letters. Part B, Applications : an international journal of research and surveys, vol. 6, no. 11, pp. 3095\u2013 3101, 2015. [2] A. El Hajjaji and M. Ouladsine, \u201cModeling and nonlinear control of magnetic levitation systems,\u201d IEEE Transactions on industrial Electron- ics, vol. 48, no. 4, pp. 831\u2013838, 2001. [3] W. Barie and J. Chiasson, \u201cLinear and nonlinear state-space controllers for magnetic levitation,\u201d International Journal of systems science, vol. 27, no. 11, pp. 1153\u20131163, 1996. [4] D. L. Trumper, S. M. Olson, and P. K. Subrahmanyan, \u201cLinearizing control of magnetic suspension systems,\u201d Control Systems Technology, IEEE Transactions on, vol. 5, no. 4, pp. 427\u2013438, 1997. [5] M. Fujita, F. Matsumura, and K. Uchida, \u201cExperiments on the h disturbance attenuation control of a magnetic suspension system,\u201d in Decision and Control, 1990., Proceedings of the 29th IEEE Conference on. IEEE, 1990, pp. 2773\u20132778. [6] C.-M. Huang, J.-Y. Yen, and M.-S. Chen, \u201cAdaptive nonlinear control of repulsive maglev suspension systems,\u201d Control Engineering Practice, vol. 8, no. 12, pp. 1357\u20131367, 2000. [7] R.-J. Wai and J.-D. Lee, \u201cRobust levitation control for linear maglev rail system using fuzzy neural network,\u201d Control Systems Technology, IEEE Transactions on, vol. 17, no. 1, pp. 4\u201314, 2009. [8] J. Xu and Y. Zhou, \u201cA nonlinear control method for the electromagnetic suspension system of the maglev train,\u201d Journal of Modern Transporta- tion, vol. 19, no. 3, pp. 176\u2013180, 2011. [9] I. U. Vadim, \u201cSurvey paper variable structure systems with sliding modes,\u201d IEEE Transactions on Automatic control, vol. 22, no. 2, pp. 212\u2013222, 1977. [10] J. Y. Hung, W. Gao, and J. C. Hung, \u201cVariable structure control: a survey,\u201d Industrial Electronics, IEEE Transactions on, vol. 40, no. 1, pp. 2\u201322, 1993. [11] N. Al-Muthairi and M. Zribi, \u201cSliding mode control of a magnetic levitation system,\u201d Mathematical Problems in Engineering, vol. 2004, no. 2, pp. 93\u2013107, 2004. [12] N. Boonsatit and C. Pukdeboon, \u201cAdaptive fast terminal sliding mode control of magnetic levitation system,\u201d Journal of Control, Automation and Electrical Systems, pp. 1\u20139, 2016." + ] + }, + { + "image_filename": "designv11_34_0002546_978-3-319-47500-4_4-Figure4.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002546_978-3-319-47500-4_4-Figure4.2-1.png", + "caption": "Fig. 4.2 Separation of boundary layer from an object leading to formation of vortices", + "texts": [ + " When a body interacts with external fluid flow a boundary layer is formed. In this layer of fluid, which is in the immediate vicinity of the body, the effects of fluid viscosity are significant. Depending on how strongly curved the body is, the flow may separate at a certain distance from the leading edge forming a wake of slow fluid behind the body. The detached boundary layer becomes a free shear layer containing vorticity that can turn into the formation of strong vortices within the wake (Fig. 4.2). Behind thick bodies, the vortices are not formed symmetrically in the wake (with respect to the midplane), forming a vortex street (i.e. two rows of vortices which are staggered with respect to each other), and causing unsteady lift forces (perpendicular to the oncoming flow) to develop, thus leading to unsteady motion transverse to the flow. This motion of the body transverse to the flow is a vibration induced by the formation of the vortices and hence the name: Vortex-Induced Vibrations. VIV can be evidently seen when a cylinder is flexibly mounted and placed in flowing water, or dragged through still water" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002456_hnicem.2014.7016234-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002456_hnicem.2014.7016234-Figure5-1.png", + "caption": "Fig. 5. Manual tool change operation", + "texts": [ + " Just like any CNC machines, the CNC router uses a computing software and breakout board or motion controller (see Fig. 4) to drive a mechanical system [4]. Then, the controller will command the motors to produce a desired toolpath from the g-codes created by the CAM software. The variable frequency drive (VFD) or inverter drive is responsible in controlling the speed of the spindle motor. Table 1 provides the technical specifications of the \u2018Super Lilok\u2019 machine. From the specification, the tool change operation currently employed is manual. Figure 5 shows the manual tool change operation using set of tools like a spanner to remove the tool from the spindle motor attachment. The next tool will be secured to the spindle attachment which is the same operation in removing the tool but in the reverse direction. In manufacturing processes, especially in metal cutting industry, full automation is possible today through the implementation of automatic tool changer (ATC) systems in the machining centers [5, 6]. The smooth run on the tool change operation sequence is defined by two (2) ATC main characteristics: the auxiliary time needed to exchange the tool or the tool-to-tool time and the number of tools it can manage in the magazine [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure15-1.png", + "caption": "Fig. 15 States of the RL2 Hand", + "texts": [ + " To make this possible there are some guides that help the cradles to slide and position each finger into the necessary configuration (Fig. 14). One of the main reasons why this design has been chosen is the better use of the free space inside the anchor. Fingers inside the anchor are in a particular position and have its joints rotating. However, the fingers keep a small chamfer in its last phalange to help it not to go out to the anchor\u2019s exterior surface. On the other hand the objective of doing free cradles is to eliminate the middle state of the hand when it is taken out from the anchor, saving time and internal mechanisms (Fig. 15). Another improvement this design presents is the adding of three strength sensors for the palm in a way that it does not only help to know the value of the strength done with it, but also its position on the design depending on the strength done in each point. 36 R. Cab\u00e1s Ormaechea These three sensors are located in the states mechanism selector, the MultiState Actuator, which interacts with the palm and is taken out to cover the ASIBOT contacts, when the Main Platform is removed. The extraction system of the RL2 Hand fingers keeps almost intact its functioning philosophy because the results obtained at the RL1 Hand has been successful" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002656_icamechs.2016.7813415-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002656_icamechs.2016.7813415-Figure2-1.png", + "caption": "Fig. 2. Mechanism of trimmer", + "texts": [ + " In this study, we researched and developed a mowing system that can mow close to obstacles. We designed a trimmer-type mowing system that can mow close to obstacles while maintaining safety [7]. We improve performance of mowing system can cut grasses and hard stem (e.q. Canada golden-rod). We then experimentally evaluated the performance of our proposed mowing system. II. SPECIFICATIONS OF TRIMMER-TYPE MOWING SYSTEM The trimmer-type mowing system presented in this work has a fixed edge and a moving edge, as shown Fig. 2. There are two advantages of using a trimmer-type mowing system. 23978-1-5090-5346-9 / 16 / $31.00 \u00a92016 IEEE If a cutter edge comes into contact with people, it can cause serious injury. In the case of the proposed system, the fixed edge of the trimmer extends beyond the moving edge. Hence, the fixed edge is the first to contact any object in the path of the system. This edge is designed so that people can be safe even after a collision. In addition, when the system is mowing close to an obstacle such as a wall, the trimmer does not damage the wall as a rotary cutter would" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure5.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure5.6-1.png", + "caption": "Fig. 5.6 Same cylinder with three different local reference frame assignments", + "texts": [ + " 4.2 Roll, pitch and yaw angles [image obtained for internet] . . . . . . 193 Fig. 4.3 Geometric interpretation for some Euler angles. Left: zxz convention. Right: zyz convention. First an a-rotation around z, so that the neutral axis N (either the new x or y axes) is properly aligned. Second a b-rotation around the N axis. Finally a -rotation around the newest z-axis so that the frame reach the final orientation. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 198 xx List of Figures Fig. 5.6 Same cylinder with three different local reference frame assignments . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 243 Fig. 5.7 Gravity effect over an free-floating rigid mass . . . . . . . . . . . . . . 262 Fig. 6.1 Linear and angular velocities associated to geometric origin g of frame R1 and linear and angular velocities of the center of mass cm . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 278 Fig. 8.3 Left: Swedish wheel, also known asMecanum wheel, invented by the Swedish engineer Bengt Ilon", + ": [a\u00d7]2 = [aaT ]\u2212 \u2016a\u20162I , the inertia tensor can also be written as, (Kane and Levinson 1985): Ic = \u222b B [\u2016r\u20162I \u2212 [rrT ]] dm \u2022 Notice that the elements of the inertia tensor depend not only in the mass of the body, but also in its geometric distribution around its center of mass. 5.2.1.1 Different Coordinates Expressions of the Inertia Tensor Consider there are multiple ways to describe the Inertia Moment Matrix of a rigid body, related to the choice of the local reference frame as for example in the case of the cylinder in Fig. 5.6. Then it arise different Inertial Moment Matrices: I(i) c = \u2212 \u222b B [ r(i)\u00d7]2dm I(j) c = \u2212 \u222b B [ r(j)\u00d7]2dm The equivalence between these different expressions arise from the coordinates transformation (3.25) between the different references: 5.2 The Kinetic Energy 243 r(i) = Rj ir (j) \u21d4 r(j) = Rj i T r(i) such that (5.24) yields I(i) c = \u2212 \u222b B [( Rj ir (j) ) \u00d7 ]2 dm = \u2212 \u222b B [( Rj ir (j) ) \u00d7 ][( Rj ir (j) ) \u00d7 ] dm which after property (1.99g) (i.e.: [(Ra)\u00d7] = R[a\u00d7]RT ) it becomes I(i) c = \u2212 \u222b B Rj i [ r(j)\u00d7]Rj i T Rj i [ r(j)\u00d7]Rj i T dm = \u2212Rj i \u222b B [ r(j)\u00d7][r(j)\u00d7]dm Rj i T meaning: I(i) c = Rj iI (j) c Rj i T \u21d4 I(j) c = Rj i T I(i) c Rj i (5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000844_icarcv.2014.7064568-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000844_icarcv.2014.7064568-Figure4-1.png", + "caption": "Fig. 4. Coordinate frame assignment to the right side of the rover.", + "texts": [ + " At any time t, the rover has an instantaneous coordinate frame R attached to the base (center point of the rover on the terrain) of the rover with respect to its global coordinate frame G, as shown in Fig. 3. Let the rover configuration vector be H=[X , Y , Z, \u03d5x, \u03d5y , \u03d5z] T defined with respect to G, where (X,Y, Z) are the coordinates of the rover and (\u03d5x, \u03d5y, \u03d5z) are the roll, pitch and yaw angles of the rover, respectively. Each key point of the rover has been assigned a coordinate frame, as shown in Fig. 4. The relations between frames attached to the links (i\u22121) and i are derived using the homogeneous coordinate transformation matrix, which consists of the Denavit-Hertenberg (D-H) parameters [18], given by the following matrix: Ti\u22121,i = \u23a1 \u23a2\u23a2\u23a3 c\u03b6i \u2212s\u03b6ic\u03b4i s\u03b6ic\u03b4i aic\u03b6i s\u03b6i c\u03b6ic\u03b4i \u2212c\u03b6is\u03b4i ais\u03b6i 0 s\u03b4i c\u03b4i di 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6 . (2) It is considered here that, at any time t, each wheel has only one contact point with the surface and the wheel is a circular rigid reference. Another assumption is that the rover wheel will not slip on the terrain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001090_icems.2014.7014015-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001090_icems.2014.7014015-Figure2-1.png", + "caption": "Fig. 2. Synchronous operation of the DSBDFIG", + "texts": [ + " Operational Principle In DSBDFIG, there are two electrical ports: the outer stator windings (with po pole pairs) and the inner stator windings (with pi pole pairs). The specialized rotor is shorted itself, thus no slip rings or brushes are needed. The outer stator and the inner stator are coupled via the rotor, of which the equivalent number of pole pairs is selected as pr=po+pi. The DSBDFIG is expected to operate in the doubly-fed synchronous mode, in which the rotor field induced by outer stator excitation synchronizes with the inner stator field, and vice versa. Fig. 2 shows the speed relationship between stator and rotor fields for synchronous operation, where o and i are electrical angular frequencies of the two stator fields, ro= o\u2013po m and ri= i\u2013 pi m are electrical angular frequencies of the rotor field induced by outer stator and inner stator respectively, m is the shaft angular frequency of the rotor. The synchronous condition is established if the two induced rotor frequencies ro and ri are identical (that is ro=\u2013 ri). Then the shaft angular speed of DSBDFIG can be expressed as io io m pp + += \u03c9\u03c9\u03c9 (1) When i=0, the DSBDFIG operates at natural-synchronous speed, the condition when i is positive or negative will be referred to as super-synchronous mode and sub-synchronous mode respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002006_s11071-016-2788-z-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002006_s11071-016-2788-z-Figure7-1.png", + "caption": "Fig. 7 Approximation of the elastic blade", + "texts": [ + " (24) Substitution of (1) into the above equations implies that \u03b2 f c = \u03b4c \u00b7 l\u03b4l2 l\u03b2 f l1 , \u03b2 f s = \u03b4s \u00b7 l\u03b4l2 l\u03b2 f l1 , (25) which are the relations between cyclic pitches and flapping angles of the stabilizing bar in steady states. 3.2 Flapping dynamics of main rotor blades Although no flapping hinges are installed to the main rotor blades, the flapping dynamics does exist due to the elasticity of the blades. For simplicity, the elastic blade is approximated into a rigid blade with a hinge spring, as is illustrated in Fig. 7. Similar to that of the Bell\u2013Hiller stabilizing bar, the equilibrium equation for the blade is given by Ib(\u03b2\u0308 + \u03a92\u03b2) = Ib (\u03b3b 8 \u03a92\u03b8 \u2212 \u03b3b 8 \u03a9\u03b2\u0307 ) \u2212 R3 \u2212 R3 0 6 \u03c1ac\u03a9vi \u2212 k\u03b2\u03b2 (26) where Ib = \u222b R R0 mbr2dr denotes the moment of inertia of the blades; \u03b3b \u03c1ac(R4 \u2212 R4 0)/Ib; k\u03b2 is the equivalent spring coefficient; and k\u03b2\u03b2 represents the torque exerted on blade by the hinge spring. Flapping dynamics of the main rotor blades can be obtained by substituting (9) into the above equilibrium equation and equating the same harmonics: \u03b2\u03080 + \u03b3b 8 \u03a9\u03b2\u03070 + (\u03a92 + k\u03b2 Ib )\u03b20 = \u03b3b 8 \u03a92\u03b80 \u2212 R3 \u2212 R3 0 6Ib \u03c1ac\u03a9vi , \u03b2\u0308c+ \u03b3b 8 \u03a9\u03b2\u0307c+2\u03a9\u03b2\u0307s+ \u03b3b 8 \u03a92\u03b2s + k\u03b2 Ib \u03b2c = \u03b3b 8 \u03a92\u03b8c, \u03b2\u0308s+ \u03b3b 8 \u03a9\u03b2\u0307s\u22122\u03a9\u03b2\u0307c\u2212 \u03b3b 8 \u03a92\u03b2c + k\u03b2 Ib \u03b2s = \u03b3b 8 \u03a92\u03b8s, (27) where \u03b20 denotes the zero-order harmonics (coning angle); \u03b2c and \u03b2s are the first-order cos and sin harmonics, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002879_s10878-020-00533-z-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002879_s10878-020-00533-z-Figure1-1.png", + "caption": "Fig. 1 a Two robots u and u\u2032 at the link position between close trajectories. b Shifting operation between trajectories with the same travel direction. c Shifting operation between trajectories with opposite travel directions", + "texts": [ + " The authors study an abstraction of the problem on a simpler scenario where the trajectories are pairwise disjoint circles of unit radius, a tour of a robot in a circle takes one time unit and the communication range of each robot is a value \u03b5 < 0.5. We say that two circles Ci and C j are close if the distance between their centers is less than 2+ \u03b5 and we denote the segment connecting their centers by {i, j}. When two robots u and u\u2032 traversing Ci and C j , respectively, lie on {i, j} at the same time, they can exchange information (they are within communication range of each other). In this case, we say that u and u\u2032 are in link position (see Fig. 1a). Moreover, since the robots travel along their trajectories at the same constant speed of 2\u03c0 length per time unit, they are synchronized and can exchange information once per time unit. Suppose now that we have set up a system with one robot per circular trajectory such that every pair of robots in close circles are synchronized. The performance of this system is compromised when a robot abandons the mission to refuel or due to some technical failure. D\u00edaz-B\u00e1\u00f1ez et al. (2015) and D\u00edaz-B\u00e1\u00f1ez et al. (2017) also address this problem and propose the following strategy: let be a link between close trajectories Ci and C j with synchronized robots u and u\u2032, respectively. Suppose that u\u2032 in C j abandons the system. When u in Ci arrives at , its communication interface detects the absence of u\u2032. To compensate, u assumes the subtask assigned to u\u2032 in C j (which is interrupted at themoment), leavingCi and shifting toC j (see Fig. 1b, c). This approach is referred as the shifting strategy in D\u00edaz-B\u00e1\u00f1ez et al. (2017). Notice that, due to kinematic constraints while shifting, it is convenient that synchronized robots fly in opposite directions, one in clockwise and the other one in counterclockwise direction (see Fig. 1). Let T = {C1, . . . ,Cn} be set of unit circles (trajectories) such that Ci and C j are disjoint for all i = j . Let \u03b5 < 0.5 be the communication range of the robots. Let G\u03b5(T ) = (V , E\u03b5) be a graph whose nodes are the centers of the circles in T and whose edges are the pairs of circle centers at distance 2 + \u03b5 or less. From now on assume that G\u03b5(T ) is connected, as otherwise there cannot be full communication in the system. As proposed in D\u00edaz-B\u00e1\u00f1ez et al. (2015) and D\u00edaz-B\u00e1\u00f1ez et al. (2017), a synchronized communication system (SCS) is a bipartite connected graphG = (V , E), a spanning subgraph of G\u03b5(T ), such that {i, j} \u2208 E if and only if the robots in Ci and C j are synchronized and fly in opposite directions", + " Since the X -motion of each of these tokens is periodic and has the same constant speed, it is easy to see that after some time, say t units of time, z and z\u2032 will meet each other in the X -axis projection. Thus, the tokens z and z\u2032 have the same x and y coordinates in the movement lattice at time t + t . Since z and z\u2032 at time t + t belong to two different robots (one robot is u and the other robot is not necessarily v), these robots are at the link position between two neighboring trajectories as shown in Fig. 1a. Therefore, robot u is not starving, a contradiction. (\u21d0) It suffices to show that, if a robot u is not starving, then there is another robot in the same row or column at some time. Robot u will be communicating with another robot, say v, at some time t . There are two possibilities for the link position in the grid. The trajectories of u and v at time t are either in the same row or in the same column. Either way, the theorem follows. Finally, by induction on k and the pigeonhole principle, we conclude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002990_012029-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002990_012029-Figure1-1.png", + "caption": "Figure 1. Gear pair kinematics.", + "texts": [ + " In the present work the model of the gear pair for three dimensional nonlinear dynamics analysis of rotary machines has been developed. The model of the gear pair is aimed to analyse a wide spectra of gear types: spur gears, helical gears, conical gears and internal gears, and takes into account such effects as variable gear mesh stiffness, gear mesh damping, transmission errors and backlash. 2. Basic definitions and state vector The element formed by two geared wheels with centers A and B is considered at Fig.1. Initial position of each wheel center is given by and . Two triad of unit orthogonal vectors are defined at each wheel center: , , is attached to the first wheel, with origin at wheel centre \u201cA\u201d, and , , is attached to the second wheel, with origin at wheel centre \u201cB\u201d. Both triads have their first vector oriented towards the contact point between wheels, their third vector perpendicularly oriented to the wheel plane (figure 1) and their second vector define each triad as right-handed coordinate system. rA0 rB0 eA10 e A 20 e A 30 eB10 eB20 eB30 ToPME-2019 IOP Conf. Series: Materials Science and Engineering 747 (2020) 012029 IOP Publishing doi:10.1088/1757-899X/747/1/012029 The position of each wheel center at current configuration is and are the translation vectors of wheel centers. The orientation of both triads at the current configuration is obtained applying the rotation tensor to each wheel to the vectors at the initial configuration: and the rotation matrices at nodes A, B, which are related to the rotation vectors and [2, 3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002120_s10556-015-0060-x-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002120_s10556-015-0060-x-Figure1-1.png", + "caption": "Fig. 1. Calculation of the discharge process.", + "texts": [ + " In the case of a suction process, the fi rst section should be selected directly in front of the suction valve while the second section is assumed to coincide with the piston crown. Let us consider a piston pump as a subject for analysis. For the discharge process, we assume that the fi rst section coincides with the piston crown and select the second section to be located immediately behind the discharge valve. Since the processes of suction and discharge are extremely similar in terms of physical sense, to make the discussion concrete we will consider the discharge process (cf. Fig. 1). Without signifi cant loss of precision of the model, in integrating the Bernoulli equation we will assume that the fl uid is incompressible. Then we may write Eq. (8) in the form: z1 + p1 \u03c1g + w1 2 2g = z2 + p2 \u03c1g + w2 2 2g + lt1\u22122 g + lf1\u22122 \u03a6 g , (15) where \u03c1 is the density of the working medium. From (15), it follows that p1 \u03c1g = p2 \u03c1g + (z2 \u2212 z1)+ w2 2 \u2212w1 2 2g + lt1\u22122 g + lf1\u22122 \u03a6 g . (16) Let us now fi nd the values of the quantities that occur on the right-hand side of (16). The difference (z2 \u2013 z1) is de- termined thus: z2 \u2212 z1 = (Sd" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000302_978-1-4471-4141-9_29-Figure29.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000302_978-1-4471-4141-9_29-Figure29.2-1.png", + "caption": "Fig. 29.2 SwarmItFIX mobile base and Bench pin module", + "texts": [ + " The workbench provides a suitable surface in which to move and securely maintain the fixture in the desired location, air and electrical supply for the whole system is embedded in it. The bench is a planar design with 52 pin modules arranged in interspersed columns in which the agent can dock. To be able to sustain the harsh manufacturing environment filled with swarf and cooling fluids, the workbench implements a cleaning system in which from the locking pin modules the chips are blown, allowing a non-stop operation for longer time periods and without human intervention. The exploded view of the pin module is shown in Fig. 29.2a. The installed bench can be appreciated in Fig. 29.1. The bench and hence the whole system can be mounted in different positions; Table 29.1 Estimate of main savings offered by one SwarmItFIX installation in the aerospace manufacturing sector Aerospace sector Automotive sector Savings Savings Conventional fixturing 2,000,000 C= Conventional fixturing 840,000 C= Fixture storage and management 100,000 C=/year 9 5 years Fixture storage and management 115,000 C= /year 9 5 years 95% Reduction of set-up times 720,000 C=/year 9 5 years 75% Reduction of setup times 600,000 C=/year 9 5 years Lower time cycle, considered 975 h/ year saved 78,000 C=/year 9 5 years Less re-positioning of part, less manufacturing operations to adapt the support device to the part local shape, considered 1200 h/ year saved 98,500 C=/year 9 5 years Total saving during the first 5 years 6,490,000 C= Total saving during the first 5 years 4,907,500 C= Fig", + "1 Final installation of the SwarmItFIX fixture system prototype horizontally (floor), vertically (walls) or even upside-down (ceiling), increasing the modularity and customization of the fixture system. The mobile base is the support of the PM, hosting all the electrical components necessary for the communication and control. It gives the first position tuning of the system, using the swing locomotion method [20]; and a 3 legged design, with modified commercial Schunk lockers for the docking mechanism, described in depth in [10], the exploded view is shown in Fig. 29.2, letting appreciate all the components, for the physical prototype see Fig. 29.1. The three legs are inserted alternatively in the bench taking electrical power and air supply from it and transmitting it to the whole agent. The base rotates about one locked leg while the other two are inside the base; it rotates at an angle and stops with the unused legs above free bench pins. Then, the latter legs slides down one at a time, two legs are again pulled in leaving one clamped. The presence of one leg always inserted in the bench guarantees the electrical and air supply; and the special gear transmission provides the necessary torque for the agent to rotate orienting at the same time the retracted legs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003181_tase.2020.2993277-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003181_tase.2020.2993277-Figure1-1.png", + "caption": "Fig. 1. 3-D structure of LDNSA.", + "texts": [ + " The stability of the impedance control scheme is discussed based on the proposed improved Bouc\u2013Wen hysterical compensation. In the following content, the principle of nonlinear stiffness and structure of the LDNSA are introduced in brief in Second II. Then, a state-space analysis based on a dynamic model as well as the improved Bouc\u2013Wen model and impedance control are applied to describe an LDNSA, followed by a stability analysis in Section III. Section IV illustrates the experimental prototype and the effects of the improved compensation algorithm. Fig. 1 shows that the LDNSA is mainly composed of a support frame, a motor-gearbox-encoder combination, a pulley with screw thread, an inner cylinder, an outer cylinder, wires, a magnetic linear encoder wrapped on the surface of the inner cylinder, three rollers, and three elastic components. The pulley is fixed to the shaft of the gearbox, and the outer cylinder is connected with the pulley by wires, which eliminates backlash. The roller is fixed to the outer cylinder with a flange shaft. Three elastic components are installed evenly on the surface of the inner cylinder with fastening screws" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001049_ijat.2013.053164-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001049_ijat.2013.053164-Figure4-1.png", + "caption": "Figure 4 Drill resharpening sequence", + "texts": [ + " The fixture was mounted to a universal tool grinder with a single-speed work spindle and fixed grinding axis. Cup-type, vitrified diamond grinding wheel was used for material removal with the intention to maintain the flatness of the cutting edges. The drill was mounted to the gun drill fixture with the secure of the carbide tip on a V-clamp and the support of the long drill stem with an extended balancing rest. The main geometries to be regenerated in sequence were the primary relief, the secondary relief, the inner relief, the oil clearance, the front clearance and the lead-in chamfer as shown in Figure 4. Each geometry was generated through manual manipulation of the swing-, tilt- and twist-axis in order to yield the right combination of compound angles. After the main axes were set, the drill was then fed into the rotating grinding wheel to remove damage areas on different parts of the drill primarily, the cutting edges and bearing pads. By the end of the six-step resharpening cycle, fresh cutting edges and geometries were regenerated. All the new and resharpened drills were subjected to rapid qualitative and quantitative assessments with a tilt-able digital microscope on a customised setup as shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003773_tie.2020.3038085-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003773_tie.2020.3038085-Figure1-1.png", + "caption": "Fig. 1. Multi-modes flow control valve (MMFCV). (a) Prototype of the MMFCV. Subscript \u201c3dp\u201d refers to the 3D printed part. (b) The channel is pressed by the pressing block. (c) The channel returns to its natural length.", + "texts": [ + " Authorized licensed use limited to: UNIVERSITY OF NEW MEXICO. Downloaded on May 17,2021 at 02:04:18 UTC from IEEE Xplore. Restrictions apply. 0278-0046 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. The MMFCV worked through the sealing and deformation of the latex membrane, the elasticity of the rubber band, and the closure of a soft silicone tube by radial pressure. As shown in Fig. 1, this was divided into the upper and lower chambers by the membrane, and these were connected to two ports denoted as N (negative)-port and P (positive)-port, respectively. The soft tube channel was passed through the lower chamber, and the ends were connected by the other two ports denoted as I (input)-port and O (output)-port. A pressing block fixed under the membrane was pulled by the rubber band. When the four ports were connected to the atmosphere, under the pulling force of the rubber band, the pressing block pressed the soft tube along the radial direction", + " Actuator II does not rapidly lose the air pressure, and can hold a part of the air at the steady state. (iii) Double-port control mode (DPC mode) (Fig. 3(d)): three actuators are respectively connected to the N -port, P - port and I-port of the MMFCV, and the O-port is connected to the atmosphere. According to the pressure difference between actuator I and actuator II, the channel can be switched between the \u201copen\u201d state and \u201cclosed\u201d state to regulate the inflation and deflation of actuator III. As shown in Fig. 1(a), the MMFCV comprised of a bottom shell, an upper shell, a pressing block, a T-connector, a ring shim, a soft tube, and a hemispherical membrane. The main body was fabricated by 3D printing technology. Furthermore, the shells were fabricated by transparent materials to easily observe the motion mechanism of the MMFCV, and the pressing block and T-connector were made of ABS. The fabrication steps are, as follows (Fig. 4): Step I: a 35 mm long soft tube was passed through the I-port and O-port of the bottom shell, and the contact edges were sealed by adhesive", + " Step III: the T-connector was inserted through the hole drilled in the center of the hemispherical membrane, and the ring shim was used to clamp the membrane with the assembly from step II. Then, these were fixed by adhesive. Step IV: the ends of the rubber band of the assembly from step III were inserted into the two holes of the bottom shell of the assembly from step I. Then, the upper shell clamps the edge of the membrane of the assembly from step III with the assembly from step I. The contact edges were sealed by adhesive. The prototype is shown in Fig. 1(a), and the specifications are listed in Table I. To quantitatively analyze the performance of the MMFCV in the three work modes, the pressure characteristics within inflation and deflation were modeled in this section. The force analysis was performed in a plane determined by the tube and rubber band. The coordinate system was established, as shown in Fig. 5(a). The origin O was fixed at the center of the bottom surface of the MMFCV, and the X-axis and Y -axis were along the radial and axial directions of the MMFCV, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure11-1.png", + "caption": "Figure 11. Tooth contact with no errors of alignment and errors of alignment q\u00bc 0.1 mm; (a) the contact path with no errors of alignment and q\u00bc 0.1 mm; (b) the contact pattern on curve-face gear and (c) the function of transmission errors.", + "texts": [], + "surrounding_texts": [ + "Based on the mathematical equations (19) of L \u00f0 f \u00de 12 and (24) of L \u00f0 f \u00de 13 , the contact point with no errors of alignment in the meshing process of curve-face gear pair can be represented as follows L \u00f0 f \u00de 13 \u00bc L \u00f0 f \u00de 12 L \u00f0 f \u00de 12 \u00bc 0 L \u00f0 f \u00de 13 \u00bc 0 8>>< >>: \u00f025\u00de So, the parameters ( 2, , , l) of contact point about rotating angle 1 of non-circular gear can be obtained from equation (2) in the meshing process of curve-face gear pair. With the mathematical equation (10) of tooth surface 3 in coordinate O3 X3Y3Z3, the contact point on curve-face gear can be established. And the contact point on noncircular gear can be established with the mathematical equation (7) of tooth surface 2 in coordinate O2 X2Y2Z2. at The University of Melbourne Libraries on June 5, 2016pic.sagepub.comDownloaded from" + ] + }, + { + "image_filename": "designv11_34_0000677_aeit.2014.7002051-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000677_aeit.2014.7002051-Figure12-1.png", + "caption": "Fig. 12: Schematic representation of the Hely prototype used as a drilling machine.", + "texts": [ + " Finally, from 42 to 72 seconds the combination of the two movements is set, imposing an helical motion to the applied load. The comparison between the output speed of both rotary and tubular machine and the reference speed required from the application are shown in figures 10 and 11. As for the first simulation, the speed trends of both linear and rotary machines are almost conicident with the reference speed, validating the proposed control system. In a third simulation the Hely prototype is used as a drilling machine, as shown in fig. 12. In this case, in the first 21 seconds of the simulation a linear movement is applied by driving LI with a reference speed of 1 m/s, while from 25 seconds to 45 seconds a rotary movement is imposed by driving RI with a reference speed of 90 rad/s. Finally, from 51 to 81 seconds, both Primaries are driven and, therefore, an helical motion to the applied load is imposed. The simulation results are shown in figures 13 and 14. As for the previous simulations, there\u2019s good agreement between the speed trends of both linear and rotary machines, validating the proposed control system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003248_s00170-020-05597-z-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003248_s00170-020-05597-z-Figure5-1.png", + "caption": "Fig. 5 The guide joints: a ball type; b roller type", + "texts": [ + " Note two details: (1) i is the number of load rings, which means that for the nuts without preload, it is the total number of rings (the product of the number of loops and the number of cycles contained in each loop), and for the pre-tightened nut from the middle position, it is half the total number of rings. Many companies follow this principle for labeling. (2) The difference between the double-nut ball screw joint and the single-nut ball screw joint is the pre-tightening method, which means that the stiffness of the double-nut and single-nut ball screw joints with the same parameters is the same. Therefore, the single-nut ball screw joint can also be calculated by Eq. (8). For the ball-type guide joint as shown in Fig. 5a, the four principal curvatures can be written as \u03c111 \u00bc \u03c112 \u00bc 2=db \u03c121 \u00bc \u22122=krdb \u03c122 \u00bc 0 9= ; \u00f09\u00de The equilibrium and compatibility conditions of the guide joint can be written as Ng 2 P1\u2212P2\u00f0 \u00desin\u03b2 \u00bc Fz P2=3 1 \u2212P2=3 0 \u00bc P2=3 0 \u2212P2=3 2 \u03b4z \u00bc 2\u03b4g=sin\u03b2 9>= >>; \u00f010\u00de where Ng is the number of balls in contact with the guide in a single slider, P1 and P2 are the forces on the upper and lower raceways under the normal load Fz, respectively, \u03b4z is the normal displacement of the slider, and \u03b4g is the deformation of half ball under the loadFz. Similar to the ball screw joint, by using Eqs. (2) and (10), stiffness of the ball-type guide joint can be calculated by Kgn \u00bc 3 8 Ncos2\u03b2C\u22121 h P1=3 1 \u00fe P1=3 2 Kgt \u00bc 3 8 Nsin2\u03b2C\u22121 h P1=3 1 \u00fe P1=3 2 9>= >;;P2\u22650 \u00f011\u00de For the roller-type guide joint as shown in Fig. 5b, by replacing Eq. (2) with Palmgren formula, it can be written as \u03b4 \u00bc CpP9=10 Cp \u00bc 3:81 l4=5 2 1\u2212\u03bd2\u00f0 \u00de \u03c0E 9=10 9>= >; \u00f012\u00de where l is the length of the roller. Derivation process similar to the ball-type guide joint, stiffness of the roller-type guide joint can be calculated by Kgn \u00bc 5 18 Ncos2\u03b2C\u22121 p P1=10 1 \u00fe P1=10 2 Kgt \u00bc 5 18 Nsin2\u03b2C\u22121 p P1=10 1 \u00fe P1=10 2 9>= >;;P2\u22650 \u00f013\u00de Based on multi-scale fractal contact theory and island area distribution function [24], the load, normal contact stiffness, and tangential contact stiffness on an interface can be calculated by Pn \u00bc \u2211 k\u00bc1 kp \u222ba 0 \u03b4k a0 \u03b4 k\u22121\u00f0 \u00de p 0 pk a 0 n a 0 da 0 \u00fe \u2211 ke\u22121 k\u00bckp\u00fe1 \u222ba 0 \u03b4k a0 \u03b4 k\u22121\u00f0 \u00de p 0 epk a 0 n a 0 da 0\u00fe \u2211 kmax\u22121 k\u00bcke \u222ba 0 \u03b4k a0 \u03b4 k\u22121\u00f0 \u00de p 0 ek a 0 n a 0 da 0 \u00fe \u222ba 0 max a0 \u03b4 kmax\u22121\u00f0 \u00de p 0 ek a 0 n a 0 da 0 Kn \u00bc \u2211 ke\u22121 k\u00bckp\u00fe1 \u222ba 0 \u03b4k a0 \u03b4 k\u22121\u00f0 \u00de kepnk a 0 n a 0 da 0 \u00fe ffiffiffiffiffiffi 2\u03c0 p ED \u03c0 1\u2212D\u00f0 \u00de a 0D=2 max a 0 1\u2212D\u00f0 \u00de=2 max \u2212a 0 1\u2212D\u00f0 \u00de=2 \u03b4 ke\u22121\u00f0 \u00de K t \u00bc 2 1\u2212\u03bd\u00f0 \u00de 2\u2212\u03bd Kn 9>>>>>>>>= >>>>>>>; \u00f014\u00de where k is an index corresponding to the scale of an asperity, ke and kp represent critical scales corresponding to elastic deformation and plastic deformation, respectively, and kmax is the scale of the largest contact spot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003531_ilt-09-2019-0378-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003531_ilt-09-2019-0378-Figure1-1.png", + "caption": "Figure 1 Schematic of surface textured bearing", + "texts": [ + " The present work develops an analytical solution of Reynolds equation for evaluating the characteristics of FLTB using the method of separation of variables and the mean of eigenvalue method. The effects of aspect ratio, eccentricity ratio and boundary conditions on the characteristics of FLTB are then investigated subsequently. The correctness and accuracy of the analytical solution are verified by comparing with CFD simulations and experimental data. Finally, the stiffness coefficients and damping coefficients and the axis trajectory and stability are investigated with the help of the analytical solution. Figure 1 is the schematic and the coordinate system of FLTB. The classical governing equation of fluid lubrication is Reynolds equation, which is hard to be solved analytically. Generally, simplification is necessary for analysis. The Reynolds equation suitable for the FLTB can be written as: @ @x h3 m @p @x ! 1 @ @z h3 m @p @z ! \u00bc 6U @h @x 12V0 (1) where the first term on the right hand side represents the hydrodynamic physical wedge effect and the second term indicates squeeze effect. The derivation of the Reynolds equation suitable for the FLTB is presented in equations (A1-1) to (A1\u20134) in Appendix 1", + "01 (heavily-loaded bearing) to 10 (lightly-loaded bearing) selected for simulation. Direct stiffness coefficient Kxx in x-direction is slightly higher than that Kyy in y-direction, when Sommerfeld number S< 1. As bearing gradually shifts to high speed and light load, Kyy increases slowly while Kxx decreases simultaneously and eventually Kyy is larger than Kxx. Whether aspect ratio is 0.5 or 1, Kxx decreases with Sommerfeld number S and Kyy increases slowly and insignificantly. The reason is that in Figure 1, attitude angle w increases gradually during operation and axial position varies from O1 to O2. The axial position gradually approaches y-axis and moves away from x-axis, so, Kxx decreases gradually and Kyy increases slowly accordingly. Characteristics of surface textured bearing YazhouMao et al. Industrial Lubrication and Tribology Damping coefficients Cxx and Cyy show a similar change trend of decreasing first and then increasing. The damping coefficient Cxx is much larger thanCyywith Sommerfeld number S< 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001186_s12541-013-0101-3-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001186_s12541-013-0101-3-Figure3-1.png", + "caption": "Fig. 3 Experimental set-up in permeability test of porous bushing", + "texts": [ + " In contrast, after the laser irradiation process, whole surface of disc is covered with the re-solidified layer consisting of a great number of circular spots generated by repeating laser pulses. And there are sparsely distributing microscopic pores which would act as gas feeding holes on the bushing surface. The pores are decreased as scanning pitch DL becomes small, which would reduce permeability of porous disc. And under the condition in this study, no pore is found when the pitch DL is 0.2 mm as shown in Fig. 2(d). The permeability of porous metal discs is investigated with an experimental set-up shown in Fig. 3. The pressurized air is supplied to the flanges containing the porous disc, and the air runs through the disc from unmodified surface to the modified one. The volumetric flow rate of air q is measured for each disc by a flowmeter and the supply Table 1 irradiating conditions in surface modification of porous bushings Averaging power Q 205 (W) Pulse frequency f 400 (s-1) Duty ratio 80 (%) Diameter of spot d 1 (mm) Scanning velocity V 3 (m/min) Scanning pitch \u0394L 0.20, 0.30, 0.35 (mm) Flow rate of shield gas (Ar) 0", + " In this test, a load cell is mounted on the top of shaft to measure the axial force generated by the supply pressure of air. The set-up is placed between the table and spindle head of machining center, so that the floating height of shaft is decided by the position of spindle head. Thus, the load capacity for different gaps is investigated, and the stiffness is estimated by the response to the change of gap. Figure 6 shows the relationship between the differential pressure \u0394P (= P1 - P2) and the air flow through the porous bushings, which is obtained by the permeability test (Fig. 3) for three kinds of bushings. Though the bushing irradiated with scanning pitch of 0.2 mm was also employed, the flow rate was lower than measurable level. As can be seen, the flow rate increases linearly with the differential pressure for any bushings. This is consistent with Darcy\u2019s law13 which describes the flow of a fluid through a porous medium as follows: \u03bc t q = \u03ba S \u0394P (1) where \u03bc is the viscosity of fluid, t is the thickness of medium, S is the cross-sectional area to flow (i.e. q/S represents the flow velocity through the medium), and \u03ba is the permeability of the medium" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001395_iecon.2013.6699572-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001395_iecon.2013.6699572-Figure1-1.png", + "caption": "Fig. 1 The diagram of \u03b1-\u03b2 transformation", + "texts": [ + " Substitute \u03b4= \u03c0/6 into (3), the La\u03b4, Lb\u03b4 and Lc\u03b4 can be calculated by \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa\u23aa \u23aa \u23aa \u23a8 \u23a7 \u2212= \u2212= \u2212= )( 2 3 )( 2 3 )( 2 3 bcc abb caa LLL LLL LLL \u03b4 \u03b4 \u03b4 (6) In the a\u03b4-b\u03b4-c\u03b4 coordinate system, define the normalised vectors of La\u03b4, Lb\u03b4 and Lc\u03b4 as \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 \u23aa\u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23a7 ++ = ++ = ++ = ])()()[( 3 2 ])()()[( 3 2 ])()()[( 3 2 222 222 222 \u03b4\u03b4\u03b4 \u03b4 \u03b4 \u03b4\u03b4\u03b4 \u03b4 \u03b4 \u03b4\u03b4\u03b4 \u03b4 \u03b4 cba c c cba b b cba a a LLL L L LLL L L LLL L L (7) Thus, it is possible to model \u03b4aL using its fundamental component. Here, neglecting the higher-order harmonics, the fundamental component of \u03b4aL can be expressed as ) 6 7cos( \u03c0\u03c9\u03b4 \u2212= tLa (8) Considering the symmetric of the three phases, \u03b4aL , \u03b4bL and \u03b4cL can be represented by the unit cosine function with a fixed phase difference of 120\u00b0 electrical angle. \u23aa \u23aa \u23aa \u23a9 \u23aa \u23aa \u23aa \u23a8 \u23a7 +\u2212= \u2212\u2212= \u2212= ] 3 2) 6 7cos[( ] 3 2) 6 7cos[( ) 6 7cos( \u03c0\u03c0\u03c9 \u03c0\u03c0\u03c9 \u03c0\u03c9 \u03b4 \u03b4 \u03b4 tL tL tL c b a (9) The vector diagram of L , \u03b4L and \u03b4L is given in Fig. 1. As shown in this figure, the \u03b4aL , \u03b4bL and \u03b4cL are located in a\u03b4-b\u03b4-c\u03b4 coordinate system and the \u03b2\u03b1 \u2212 coordinate system can be represented by \u239f\u239f \u239f \u239f \u23a0 \u239e \u239c\u239c \u239c \u239c \u239d \u239b \u239f \u239f \u239f \u239f \u23a0 \u239e \u239c \u239c \u239c \u239c \u239d \u239b \u2212 \u2212\u2212 =\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u03b4 \u03b4 \u03b4 \u03b2 \u03b1 c b a L L L L L 2 3 2 30 2 1 2 11 3 2 (10) The diagram of the \u03b2\u03b1 \u2212 coordinate system is also shown in Fig. 1. As can be seen from the figure, the rotating angle \u03be can be calculated by ]arctan[ \u03b1 \u03b2\u03be L L = (11) Here, \u03be are calculated using the artan2 function, thus the numerical range of \u03be is )\u03c0,\u03c0(\u2212 . Combining (2), (8) and (11), the rotor position can be represented by \u23a5\u23a6 \u23a4 \u23a2\u23a3 \u23a1 += \u03be\u03c0 \u03c0 \u03b8 6 75.22 (12) Since the phase inductance can be identified by current slope difference method, \u03b4aL , \u03b4bL and \u03b4cL can be calculated from (6) and (7), and the rotor position can be estimated in a wide speed range. III. SIMULATION AND EXPERIMENTAL RESULTS To verify the proposed method, simulation and experiments have been done on a 12/8 structure SRM prototype" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002732_iccicct.2016.7987997-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002732_iccicct.2016.7987997-Figure1-1.png", + "caption": "Fig. 1. Motion in 6 DoF considering earth fame and body frame", + "texts": [ + " In section 3, the LMI based controller fundamentals are explained in order to design a robust depth controller. The simulation results for a REMUS AUV II model are presented in section 4 and outputs are discussed. In Section 5 results are compared for different quantitative versus qualitative controllers. UNDERWATER VEHICLE A mathematical model of a genric Autonomous Underwater Vehicle(AUV) with 6 Degree of Freedom(6-DoF) via 12 first order equations is referred in this paper. According to SNAME [9] the notation for marine vessel, the position and Euler angle components as shown in figure 1 in surge, sway,heave, roll, pitch and yaw are [x,y,z,\u03c6 ,\u03b8 ,\u03c8] and corresponding linear and angular velocity are expressed as [u,v,w, p,q,r]. The nonlinear kinematics and kinetics can be derived by defining body fixed reference frame [XB,YB,ZB] and the earth fixed reference frame [XE ,YE ,ZE ] as shown in [10] Kinematics deals with geometric aspects of motion of body, in our case AUV. The 6 Dof kinematic equations of motion can be written in component form as [10] X\u0307 = ucos(\u03c8)cos(\u03b8)+ v[cos(\u03c8)sin(\u03b8)sin(\u03c6)\u2212 sin(\u03c8)cos\u03c6 ] +w[sin(\u03c8)sin(\u03c6)+ cos(\u03c8)cos(\u03c6)sin(\u03b8)] Y\u0307 = usin(\u03c8)cos\u03b8 + v[cos(\u03c8)cos(\u03c6)+ sin(\u03c6)sin(\u03b8)sin(\u03c8)] +w[sin(\u03b8)sin(\u03c8)cos(phi)\u2212 cos(\u03c8)sin(\u03c6)] Z\u0307 =\u2212usin(\u03b8)+ vcos(\u03b8)sin(\u03c6)+wcos(\u03b8)cos(\u03c6) \u03c6\u0307 = p+qsin(\u03c6)tan(\u03b8)+ rcos(\u03c6)tan(\u03b8) \u03b8\u0307 = qcos(\u03c6)\u2212 rsin(\u03c6) \u03c8\u0307 = q sin(\u03c6) cos(\u03b8) + r cos(\u03c6) cos(\u03b8) ,\u03b8 =\u00b190o (1) The rigid-body kinetics can be derived by applying Newtonian mechanics in [10] as following first three equations represent the translational motion, while the last three equations represent the rotational motion The nomenclature is well explained by [8] for the standard equation of motion shown below, Surge, or translation along the x-axis: m [ u\u0307\u2212 vr+wq\u2212 xg(q2 + r2)+ yg(pq\u2212 r\u0307)+ zg(pr+ q\u0307) ] = XHS +Xu|u|u|u|+Xu\u0307u\u0307+Xwq +Xqqqq+Xvrvr+Xrrrr+Xprop (2) Sway, or translation along the y-axis: m [ v\u0307\u2212wp+ur\u2212 yg(r2 + p2)+ zg(qr+ p\u0307)+ xg(qp+ r\u0307) ] = YHS +Yv|v|v|v|+Yr|r|r|r|+Yv\u0307v\u0307 +Yr\u0307r\u0307+Yurur+Ywpwp+Ypq pq +Yuvuv+Yuu\u03b4r u 2\u03b4r (3) Heave, or translation along the z-axis: m [ w\u0307\u2212uq+ vp\u2212 zg(p2 +q2)+ xg(rp\u2212 q\u0307)+ yg(rq+ p\u0307) ] = ZHS +Zw|w|w|w|+Zq|q|+Zw\u0307w\u0307 +Zq\u0307q\u0307+Zuquq+Zvpvp+Zrprp +Zuwuw+Zuu\u03b4r u 2\u03b4s (4) Roll, or rotation about the x-axis: Ix p\u0307+(Iz\u2212 Iy)qr\u2212 (r\u0307+ pq)Ixz +(r2\u2212q2)Iyz +(pr\u2212 q\u0307)Ixy +m [yg(w\u0307\u2212uq+ vp)\u2212 zg(v\u0307\u2212wp+ur)] = KHS +Kp|p|p|p|+Kp\u0307 +Kprop (5) 978-1-5090-5240-0/16/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000216_gt2010-23035-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000216_gt2010-23035-Figure1-1.png", + "caption": "Fig. 1 A Beam segment and the analogous motion planes", + "texts": [ + " The rotor is considered as a free-free body in space, upon which combination of static and dynamic forces (e.g. gravity, bearing, imbalances and gyroscopic moments) can act. The rotor is modeled by circular Timoshenko beam finite elements to account for shear deformation and rotary inertia. Analogy between motion planes is constructed to relax unnecessary complexities arising from the use of a unified rotational direction convention [17]. The equations of motion of a rotating Timoshenko beam element with shear deformation included, in two perpendicular planes of motion XZ and YZ (Fig.1) can be written as follows: 2 2 2 2 ( ) ( , ) ( ) ( , ) S x z S y z x xAG p Z t m Z Z t y yAG p Z t m Z Z t \u03c1 \u03b1 \u03c1 \u03b2 \u2202 \u2202 \u2202\u23a1 \u23a4\u2212 + =\u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6 \u2202 \u2202 \u2202\u23a1 \u23a4\u2212 + =\u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6 2 2 2 2 ( ) ( ) ( ) ( ) S T r z S T r z xEI AG J J Z Z Z t t yEI AG J J Z Z Z t t \u03b1 \u03b1 \u03b2\u03c1 \u03b1 \u03c9 \u03b2 \u03b2 \u03b1\u03c1 \u03b2 \u03c9 \u2202 \u2202 \u2202 \u2202 \u2202 + \u2212 = + \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 + \u2212 = \u2212 \u2202 \u2202 \u2202 \u2202 \u2202 (1) Here the following relations hold sx x Z \u03b1 \u03c6\u2202 = + \u2202 , x sx s V G A \u03c6 \u03c1 = (2) YM EI Z \u03b1\u2202 = \u2202 (3) \u03c1 is shear form factor ( 0.75 for circular beam section, [18]). As shown in Fig. 1, the rotations \u03b1 and \u03b2 as well as shear angle in both planes are such that relation (3) is similar for both the planes. Evaluation of shape functions is achieved using the exact solution of the homogeneous form of eq. (1). Using the kinetic and potential energy expressions [26], carrying out integration over element length and applying Lagrange\u2019s equations, we get [ ] [ ] { } [ ] [ ] { } { } [ ] [ ] { } { } 1 1 1 1 1 1 0 0 0 0 0 0 00 t r e e t r y e e e y M M q q M M G k q q Q kG \u23a1 \u23a4 \u23a1 \u23a4 +\u23a2 \u23a5 \u23a2 \u23a5 \u23a3 \u23a6 \u23a3 \u23a6 \u23a1 \u23a4\u23a1 \u23a4 \u23a1 \u23a4\u23a3 \u23a6\u23a2 \u23a5+ + =\u23a2 \u23a5 \u23a2 \u23a5\u23a1 \u23a4\u2212 \u23a3 \u23a6\u23a3 \u23a6\u23a3 \u23a6 && && & (4) Here subscripts t and r refer to translation and rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003421_aim43001.2020.9158930-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003421_aim43001.2020.9158930-Figure4-1.png", + "caption": "Fig. 4. (a) 3D mesh model of furniture parts, (b) wireframe view of the object comprising triangles. Red lines represent the outline of the object.", + "texts": [ + ", Tnm }. The generation process of grasp pose candidates comprises three steps. First, a preliminary point is sampled and its approach direction is determined. Subsequently, we find the antipodal point of the preliminary point and make a grasp pose candidate using the points and the approach direction vector. Finally, the generated grasp pose is checked to see if there is any collision. The process is detailed in the following subsections. The proposed method exploits the outline of the mesh shown in Fig. 4. If the exact outline of an object is known, the optimal grasp pose of gripper can be determined based on the geometry of the gripper and object. The outline can be easily obtained in convex shape models, and there are methods that calculate the grasp pose by fitting the mesh model or point cloud into primitive shapes, such as boxes and cylinders [13]\u2013[15]. For non-convex shapes, there is a computational burden of obtaining the outline of the object (Fig. 4 (b)). We simplify this problem by considering every triangle in the mesh model. We utilize the property that the outline of the object is among all lines that form the triangles. If a point is created not on the outline but within the mesh, it can be discarded by examining whether a collision occurs between the object and the gripper that at the point. Thus, we create points on every line of the triangles and remove the points that are not approachable by the gripper. 590 Authorized licensed use limited to: Cornell University Library" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002674_smc.2016.7845000-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002674_smc.2016.7845000-Figure1-1.png", + "caption": "Fig. 1. Quadrotor Model", + "texts": [ + " The 6-DOF motion of the quadrotor is described using two right-hand reference frames; the earth inertial frame of reference and body frame of reference denoted by Eframe and B-frame, respectively. The E-frame is symbolized by (Oe, Xe, Ye, Ze) where (Oe) is the axis origin and (Xe, Ye, Ze) are in NWU configuration with respect to the earth. The B-frame attached to the body of the quadrotor and is denoted by (Ob, Xb, Yb, Zb) where (Oe) is the axis origin and coincides with the center of the quadrotor and (Xb, Yb, Zb) pointing towards the front, left and up, respectively. Figure 1 illustrates the cross structure of the quadrotor and its reference frames. The quadrotor linear position (\u03be = [x, y, z]) is determined using the vector between the origins of both frames represented in the E-frame. The system\u2019s attitude denoted by (\u0398 = [\u03c6, \u03b8, \u03c8]) is defined by the orientation of B-frame with respect to the E-frame. To transform a vector from B-frame to E-frame a rotation matrix is needed [20], and it is given by: R = c\u03b8c\u03c8 \u2212c\u03c6s\u03c8 + s\u03c6s\u03b8c\u03c8 s\u03c6s\u03c8 + c\u03c6s\u03b8c\u03c8 c\u03b8s\u03c8 c\u03c6c\u03c8 + s\u03c6s\u03b8s\u03c8 \u2212s\u03c6c\u03c8 + c\u03c6s\u03b8s\u03c8 \u2212s\u03b8 c\u03c6s\u03b8 c\u03c6c\u03b8 (1) where cx = cos(x), sx = sin(x)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure4.5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure4.5-1.png", + "caption": "Fig. 4.5 Mathematical model for flux pulsation calculation, with air-gap \u03b4, conductivity \u03c7 and permeability \u03bc, see [4.10]", + "texts": [ + " The analytical methods have been developed mainly for flux ripple on salient pole synchronous machines, but they can be applied for all other kinds of machines and for converter fed ones as well. The methods show their advantages especially at design areas with massive iron material. Here eddy-current problems often strongly dependent on the saturation of the iron. In contrast to numerical methods, which can take saturation into account, analytical methods require the magnetic properties to be linearised by using some form of equivalent permeability. In addition, an analytical method has to simplify the geometry of the real machine and use a mathematical model such as that shown in fig. 4.5. This model represents the flux pulsation by a current sheet adjacent to an airgap with a smooth surface. The \u2018pole pitch\u2019 of the current sheet is assigned with \u03bb. The magnitude of the current is adjusted to the first harmonic of the flux pulsation at the design part surface for frequency \u03c9 = 0, i.e. in the absence of eddy currents, to a level that it is the same as for the real machine. Gibbs [4.9] uses a harmonic factor to take the additional loss due to the higher harmonics into account. The influence of the mean flux density is neglected", + " The tangential field can be determined from equation 4.24 using fig. 4.7 to provide the relative permeability. The magnetization curves for the two steels, which have been considered, have been presented in the unusual form shown in fig. 4.7, because the product on the left-hand side of equation 4.24 is fixed by the operating condition and is essentially independent of the permeability. The theory in [4.8] is more accurate by the application of the two-dimensional solution of the field problem according fig. 4.5, which had been outlined by Lawrenson [4.10]. However, Lawrenson did not consider the problem of selecting a suitable permeability. Following the method in [4.9], a field-dependent permeability is introduced, but Oberretl found experimentally that better results are obtained by using only 85% of the maximum tangential field rather than the maximum value. 4.4 Eddy Current Losses within Massive Magnetic Material 55 The theory includes the damping effect of the eddy currents and makes allowance for the effect of saturation in the armature teeth tips on the higher pulsation harmonics. The losses are given by equation 4.25: ( ) L 2 el 22 \u03b4d,ri Re4 \u02c6 k fB A P S \u22c5 \u22c5 \u22c5\u22c5\u22c5 = \u03b3 \u03c7\u03bb , (4.25) where the damping effect of eddy currents has been included by modifying the original first harmonic component of the radial air-gap tooth-ripple flux density at the pole surface \u03b40, 1B\u0302 according to equation 4.26, see [4.8, 4.15], indices accord- ing fig.4.5: ( ) ( ) ( ) ( ) ( ) ( ) +\u22c5\u22c5++ \u2212\u22c5\u22c5\u2212 +\u22c5\u22c5++\u2212\u22c5\u22c5\u2212\u22c5= \u2212 \u2212 \u03b3 \u03c0 \u03bb\u03bc\u03bc\u03b3 \u03c0 \u03bb\u03bc\u03bc \u03bc\u03bc\u03bc\u03bc \u03b4 \u03bb \u03c0\u03b4 \u03bb \u03c0 \u03b4 \u03bb \u03c0\u03b4 \u03bb \u03c0 2 1 2 1 1111\u02c6\u02c6 r1 2 r3r1 2 r3 r1 2 r3r1 2 r3 \u03b40, 1 \u03b4d, 1 ee ee BB . (4.26) The skin depth parameter \u03b3 is given by equation 4.27: obr,0el 2 2 2 \u03bc\u03bc\u03c0\u03c7 \u03bb \u03c0\u03b3 \u22c5\u22c5\u22c5\u22c5\u2212 = fj , (4.27) 56 4 Additional Losses Due to Higher Voltage Harmonics The harmonic factor kL includes an allowance for the presence of tooth-ripple harmonics: 2 \u03b4m,0, \u03b40, 0 2 \u03b40, 1 \u03b4m,0, L 1 \u22c5 = \u221e = B B nB B k \u03c5 \u03bd , (4", + " The equivalent permeability of the pole face \u03bcr,Ob is determined in equation 4.29: \u03bc \u03bc \u03bb \u03c0 \u03b3 \u03bc \u03bc\u03b8 \u03b40 r,Ob d, 0 r,Ob , H B= \u22c5 \u22c5 \u22c5 \u22c52 0 85 1 . (4.29) Equation 4.29 is applied iterative in conjunction with fig. 4.7. The process is an iterative one, but converges after a few steps. Even though the analytical methods give relatively fast results, once they are implemented they should be proven by numerical methods. A finite element method [4.13] solves Helmholtz equation 4.22, and has been applied to the mathematical model of fig. 4.5. The losses have been calculated with a permeability which is constant in time. The tooth-ripple field at the pole surface has been generated by a thin current-carrying layer on the 4.4 Eddy Current Losses within Massive Magnetic Material 57 surface of the armature consisting of 10 segments in the circumferential direction, each with the same magnitude of current density but a changing phase to model relative movement with respect to the pole. This current distribution is shown in fig. 4.9. Adaptive mesh generation with automatic mesh refinement based on the flux density has been used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001473_nems.2013.6559871-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001473_nems.2013.6559871-Figure6-1.png", + "caption": "Fig. 6 The model of the beam based on graphene paper/graphene oxide paper.", + "texts": [ + " In order to analyze the actuation of actuator in detail, we adopted a classical equation (1) depicted as follows [16]: 2 222132 1221212122 222 22 111 2 212121212161 ttttttEEwwtEwtEw TttttEEww r (1) 2 221 2 12121 2 22 22 11 2 21212121 2322 61 ttttttEE TttttEE r tEtE (2) 2 221 2 121 4 2 4 1 212121 2322 61 tttttttt Ttttt r (3) Where t1, w1, 1, E1 are thickness, width, thermal expansion coefficient, and Young\u2019s modulus of graphene oxide layer, respectively. t2, w2, 2, E2 represent thickness, width, thermal expansion coefficient, and Young\u2019s modulus of graphene layer, respectively. l, r, T, are on behalf of length, radius of curvature, difference between initial temperature and final temperature of macroactuator in sequence. The details are demonstrated in Fig. 6. For a given macroactuator, l and w are the same for both graphene paper and graphene oxide paper. Therefore, equation (1) can be simplified into (2). Apparently, the curvature of the beam is proportional to T. In addition, the greater difference between 1 and 2, the greater the change of the beam. According to the literature, both of graphene paper and graphene oxide paper own a nearly equal Young\u2019s modulus, which makes the (2) be reduced into the (3). It is noteworthy that the curvature of the beam can achieve a maximum when E1 is equal to E2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001843_jahs.59.022006-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001843_jahs.59.022006-Figure1-1.png", + "caption": "Fig. 1. PAM operating principles.", + "texts": [ + " PAM actuators are simple devices that consist of four basic parts: an inner elastomeric bladder, an outer helically braided sleeve, and two end fittings that hold these components together, creating a pressuretight seal and providing a means to transfer force and deflection into the system. The operating principle of a PAM is as follows: As the bladder is pressurized with air, it expands in the radial direction and forces the braided sleeve to increase in diameter through reorientation of the helical braid. However, the fixed length of the braided sleeve filaments generates a contractile force and causes a reduction in the overall length of the muscle. Figure 1(a) shows a 2.54-cm-diameter PAM at rest and pressurized to 0.62 MPa while being allowed to freely contract. Because of this operating principle, PAMs are intrinsically a unidirectional 022006-2 actuator. While they are capable of either contraction or extension, depending on the initial braid angle \u03b80 (labeled in Fig. 1(a)), the forces created by contractile PAMs are much greater than those of an extensile PAM. For this reason, the PAMs studied here are contractile. Furthermore, contractile PAMs generate tensile forces and are, therefore, not susceptible to actuator buckling. In this configuration, PAMs behave similarly to natural muscles, as shown in Ref. 16. If controlled motion in two directions is desired, two PAMs can be arranged as an agonist/antagonist pair to produce bidirectional rotational motion about a hinge, as shown in Fig. 1(b). Here the lower PAM is the active agonist muscle, pulling on the passive antagonist upper PAM and forcing the joint to rotate clockwise. This arrangement is very similar to that of the human musculature system and is well suited for driving TEFs on helicopter blades. PAMs have been utilized for many years for prosthetic and robotic applications (Ref. 17), but only recently have they begun to be considered for aerospace applications. References 18\u201323 cover aerospace applications of PAM technology to date, where they have been considered for driving rigid TEFs (Refs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003696_s11665-020-05230-w-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003696_s11665-020-05230-w-Figure6-1.png", + "caption": "Fig. 6 The powder distribution with powder bed thickness of 64.1 lm in DEM", + "texts": [ + " When the sliced layer thickness is higher than 70 lm, the relative density of the corresponding powder bed continues to rise. Since a high relative density of the powder bed is beneficial to reduce balling (Ref 9), while too high a layer thickness might cause incomplete melting, it is necessary to find a minimum sliced layer thickness that has a relative density of the corresponding powder bed as high as possible. For 316L stainless steel powder in the simulation, the suitable sliced layer thickness is 60 lm, according to Fig. 5. The simulated powder packing results in Fig. 6 show that small particles gathered near the edge and large particles gathered in the middle of the substrate for the powder bed thickness of 64.1 lm with the relative density of 0.47. This is because small particles pass through the gap between the blade and substrate first, and large particles at the end of the substrate might pop out from the powder bed due to the squeeze between the blade and the edge of the substrate. In order to avoid the influence of the randomly packed powder bed, a small computation domain of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001770_cobep.2015.7420044-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001770_cobep.2015.7420044-Figure2-1.png", + "caption": "Fig. 2: Stator and rotor variables in dq frame.", + "texts": [ + " The 123 to odq transformation is given by: xg123= Pg gx g godq (1) in which xg is the matrix of variables (voltages, currents or flux) with g = s, r for stator and rotor variables, respectively, 978-1-4799-8779-5/15$31.00 c\u00a9 2015 IEEE and P g g is the matrix of Park transformation [20] given by: P g g = \u221a 2 3 \u23a1 \u23a2\u23a3 1\u221a 2 cos (\u03b4gg) \u2212 sin (\u03b4gg) 1\u221a 2 cos ( \u03b4gg \u2212 2\u03c0 3 ) \u2212 sin ( \u03b4gg \u2212 2\u03c0 3 ) 1\u221a 2 cos ( \u03b4gg \u2212 4\u03c0 3 ) \u2212 sin ( \u03b4gg \u2212 4\u03c0 3 ) \u23a4 \u23a5\u23a6 (2) For stator variables \u03b4gg = \u03b4g and rotor variables \u03b4gg = \u03b4g \u2212 \u03b8r, where \u03b4g is the angle between the reference frame d \u2212 axis and the stationary stator reference frame and \u03b8r is the angular position of the rotor shift. Fig. 2 shows the relationship between stator and rotor variables in dq frame and the stationary stator reference frame. Thus, for stationary stator reference frame, \u03b4g = 0, the dq variables (stator and rotor) are sinusoidal with a frequency equal to the stator current, even when the rotor speed is variable [21]. Assuming no zero sequence components, the induction generator model in the stationary reference frame is given by: vss = rsi s s + d dt \u03c6s s (3) vsr = rri s r + d dt \u03c6s r \u2212 j\u03c9r\u03c6 s r (4) \u03c6s s = lsi s s + lmisr (5) \u03c6s r = lmiss + lri s r (6) ce = 2lm (issi s\u2217 r ) (7) where xs g = xs gd+jxs gq is the vector of variable x in stationary stator reference frame, with g = s, r (s = stator and r = rotor), x = v, i, \u03c6 and and (z) is the imaginary part of z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000743_memsys.2013.6474366-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000743_memsys.2013.6474366-Figure2-1.png", + "caption": "Figure 2. Schematic of the paper-based MFC.", + "texts": [ + " Even though conventional MFCs generate a low power density, it is enough for powering the one-time use diagnostic sensors. The corresponding author has presented his glassor plastic-based micro-sized MFCs at Hilton Head 2008, 2010, and MEMS 2011 & 2012 [9-12]. Here, we described the first paper-based MFC that was capable of generating its own power when a drop of inoculum was added to the device. We also demonstrated that the paper based MFCs can be stackable in series or in parallel for a desirable operating voltage and current. As shown in Fig. 2, the paper-based MFC (4cm x 3cm) is created by sandwiching five components; (i) carbon cloth-based anode/cathode layers having oxygen plasma treatment, (ii) micro-fabricated paper anode/cathode complete the MFC operation. MEMS 2013, Taipei, Taiwan, January 20 \u2013 24, 2013978-1-4673-5655-8/13/$31.00 \u00a92013 IEEE 809 reservoirs, (iii) a paper-based proton exchange membrane (PEM), and (iv) patterned anode/cathode tapes for sample inputs. All the papers employed here were Whatman #1 filter paper. Fig 3 shows a fully assembled paper-based MFC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000760_peds.2013.6527174-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000760_peds.2013.6527174-Figure1-1.png", + "caption": "Fig. 1. Construction of segment type SRM.", + "texts": [ + " To solve the problem, we propose the novel segment type SRM which has toroidal coil windings in the stator and dual rotor structure. This paper explains about the dual rotor segment type novel SRM (DrSRM) and analyses the performance characteristics using the finite element method (FEM). On the flat type SRM, leakage flux is greatly generated at both side coil of the stator. To reduce the leakage flux we propose shield structure. Effectiveness of the shield structure is verified analytically. II. SEGMENT TYPE SRM A. Inner Rotor Segment Type SRM Fig. 1 shows construction of the inner rotor segment type SRM. Segment cores are embedded in the aluminum rotor block and the stator has full pitch three phase windings. Fig. 2 shows torque waveform of the SRM simulated by FEM. The maximum torque is twice as that of a same-sized VR SRM. The average torque is increased by about 40%. The radial force is smaller and so the vibration and acoustic noise are smaller than the VR SRM, because four poles among the six poles are always excited. The iron loss is low because the magnetic path is short [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000987_j.ijsolstr.2013.02.017-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000987_j.ijsolstr.2013.02.017-Figure3-1.png", + "caption": "Fig. 3. Evolution of the system configuration due to the increment of debonding, da > 0.", + "texts": [ + ", without movement of B) from G cb = 0 (tip position, a) to G cb > 0 (tip position, a + da) for which peeling proceeds dynamically and in uncontrolled way. We shall examine how the system evolves under the tip position increment, da. The increments of all system parameters will be evaluated in terms of da. The incremental displacement of the extremity A is: duA \u00bc dxAex \u00fe dyAey \u00f017\u00de The increments dh, de and dl are obtained in terms of da, dxA and dyA as follows. We denote by I0 the new position of the tip due to the increment da > 0 and by A0 the new position of the tape extremity, see Fig. 3. The detached length changes from AI = l to A0I0 = l + dl. The corresponding lengths in the unstretched configuration are respectively L0 and L0 + dL0. It is assumed that the width of the tape is uniform. The case of an inextensible tape with variable width will be discussed in Section 5. The pre-strain is supposed to be uniform along the interface Ox and is given by: e1 \u00bc ln\u00f0da=dL0\u00de \u00f018\u00de The portion da of attached tape, has the unstretched length dL0 and the length dl after detachment. According to Eqs. (5) and (18) we have: de \u00bc ln\u00f0\u00f0l\u00fe dl\u00de=\u00f0L0 \u00fe dL0\u00de\u00de ln\u00f0l=L0\u00de \u00bc dl l da L0 exp\u00f0 e1\u00de \u00f019\u00de From Fig. 3 it can be seen that dxA = l cos h + da (l + dl) cos (h + dh), where dxA is the incremental displacement of A in the direction x. It follows that: da \u00bc cos\u00f0h\u00dedl l sin\u00f0h\u00dedh\u00fe dxA \u00f020\u00de We have also: dyA \u00bc \u00f0l\u00fe dl\u00de sin\u00f0h\u00fe dh\u00de l sin\u00f0h\u00de \u00bc sin\u00f0h\u00dedl\u00fe l cos\u00f0h\u00dedh \u00f021\u00de With Eqs. (19)\u2013(21), de, dl and dh can be expressed in terms of da, dxA and dyA: de \u00bc sin\u00f0h\u00de dyA l \u00fe \u00f0cos\u00f0h\u00de exp\u00f0e e1\u00de\u00de da l cos\u00f0h\u00de dxA l \u00f022\u00de dl \u00bc sin\u00f0h\u00dedyA \u00fe cos\u00f0h\u00deda cos\u00f0h\u00dedxA \u00f023\u00de dh \u00bc cos\u00f0h\u00dedyA l sin\u00f0h\u00deda l \u00fe sin\u00f0h\u00dedxA l \u00f024\u00de The constraint expressed in Eq", + " Let us assume first that the debonding energy c is constant. Then, Eq. (32) takes the form: d da \u00f0G cb\u00de \u00bc F\u00f0e\u00deF 0\u00f0e\u00de lE2 1 cos h exp\u00f0e e1\u00de\u00f0 \u00de2 \u00f034\u00de It was postulated that force softening does not occur during extension of the tape, i.e., F 0 (e) > 0. If K > 0 (non-zero spring stiffness), we have d da \u00f0G cb\u00de < 0 since 0 < E2 <1 from Eq. (30). Thus, the peeling process is stable in this case. In the particular case of zero spring stiffness (K = 0), we have E2 =1 and d da \u00f0G cb\u00de \u00bc 0. Then, the critical condition G cb = 0 is maintained when the tip I (Fig. 3) is moving. Peeling proceeds without the necessity to displace B rightwards. But it should be noted that debonding would be arrested by moving B slightly to the left. K = 0 is realized when the length of the spring is infinite. Then, the tension F// is constant (i.e., does not depend on the displacement of A) and the tape is subject to the constant force, F = F//ex + F\\ey. As expected, instability is promoted when the system is softer. Let us assume now that @c @h \u00bc 0 and that @c @a < 0. This means that the debonding energy is weakening along the interface Ox and is independent of the peel angle, h", + " AM wishes to express his appreciation for the hospitality provided during his visits at Caltech. Consider a tape element with unstretched length L0. When detached, this element has the length, l and is subject to the tensile force, F. The elastic energy stored in the detached element is equal to the work of the tensile force: Wel\u00f0L0 ! l\u00de \u00bc Z l L0 F\u00f0l0\u00dedl0 \u00bc Z e 0 F\u00f0e0\u00del0de0 \u00bc L0 Z e 0 F\u00f0e0\u00de exp\u00f0e0\u00dede0 \u00f0A1\u00de Here, we have used the definition of the logarithmic strain: e = ln (l/ L0). For a given tip position a, the elastic energy stored in the parts AI and II0 (see Fig. 3) of the tape is: Wel(AI) + Wel(II0). For the position a + da this energy becomes Wel\u00f0A0I0\u00de. Thus the increment of elastic energy in the tape is: dW tape el \u00bcWel\u00f0A0I0\u00de Wel\u00f0AI\u00de Wel\u00f0II0\u00de \u00f0A2\u00de The portion AI of the tape has current length l, unstretched length L0, and current strain e = ln(l/L0), thus: Wel\u00f0AI\u00de \u00bcWel\u00f0L0 ! l\u00de \u00bc L0 Z e 0 F\u00f0e0\u00de exp\u00f0e0\u00dede0 \u00f0A3\u00de The portion II0 has current length da (see Fig. 3) and unstretched length dL0. The corresponding strain is equal to the pre-straining of the tape, e1 = ln (da/dL0). Therefore: Wel\u00f0II0\u00de \u00bcWel\u00f0dL0 ! da\u00de \u00bc dL0 Z e1 0 F\u00f0e0\u00de exp\u00f0e0\u00dede0 \u00f0A4\u00de The ligament A0I0 has current length l + dl, unstretched length L0 + dL0, and current strain e + de, then: Wel\u00f0A0I0\u00de \u00bc \u00f0L0 \u00fe dL0\u00de Z e\u00fede 0 F\u00f0e0\u00de exp\u00f0e0\u00dede0 L0 Z e 0 F\u00f0e0\u00de exp\u00f0e0\u00dede0 \u00fe dL0 Z e 0 F\u00f0e0\u00de exp\u00f0e0\u00dede0 \u00fe L0F\u00f0e\u00de exp\u00f0e\u00dede \u00f0A5\u00de Combining Eqs. (A2)\u2013(A5), we get: dWtape el \u00bc dL0 Z e e1 F\u00f0e0\u00de exp\u00f0e0\u00dede0 \u00fe L0F\u00f0e\u00de exp\u00f0e\u00dede \u00f0A6\u00de The relationship in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003558_icra40945.2020.9197469-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003558_icra40945.2020.9197469-Figure2-1.png", + "caption": "Fig. 2. Singe push to move the obstacles under the target in layer 4", + "texts": [ + " A zig-zag push is a motion where the gripper uses a zig-zag movement that contains several linear motions to push the obstacles side to side. This motion can not only move the obstacles out of the way, but also break the static contact force, such as shaking to insert a key. However, a single push is generally faster than the zig-zag push and can reduce the likelihood of the damage to the fruit. Therefore, we only use the zig-zag push in more complex situation, depending on the number and distribution of the obstacles in layer 4. 1) Single Push: As shown in Fig. 2(a), if an obstacle is located below the target (layer 4), the gripper may capture the obstacles if it moves up straightly to enclose the target. In this case, the gripper can use a single push operation to push aside the obstacle (to the right in the figure) before swallowing the target (Fig. 2(b) and (c)). Since the gripper size is limited, a single push operation makes it easier to move a few number of adjacent obstacles out of the way, but hard to separate sparsely distributed 4958 Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 04,2020 at 03:49:54 UTC from IEEE Xplore. Restrictions apply. obstacles. Therefore, ignoring the central block, we use the number of blocks nh within the largest group of adjacent unoccupied blocks (without obstacles) to determine whether to use a single push or zig-zag operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure5.8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure5.8-1.png", + "caption": "Figure 5.8 Projected views of the two closure poses for the same slider position.", + "texts": [ + " Interpreting geometrically, these visually distinct poses are such that the joint centers B+ (with \ud835\udf0e1 = + 1) and B\u2212 (with \ud835\udf0e1 = \u2212 1) happen to be the intersection points between the sphere with center C and radius CB2 = b23 and the circle with center A and radius AB1 = b12. As a special case, if r2 1 + r2 3 = h2 03, then the two poses become the same, in which \ud835\udf0301 = \ud835\udefd01 = atan2(r3, r1). In such a case, the above-mentioned sphere and circle intersect each other tangentially at a single point. The two distinct closure poses of the mechanism for the same slider position are illustrated in Figure 5.8 as projected onto the 1\u20133 plane of the base frame. On the other hand, the other closure sign variable \ud835\udf0e2 that arises in solving Eq. (5.234) for \ud835\udf1912 and \ud835\udf0312 does not lead to visually distinct closure poses of the mechanism. The fact that these poses are visually indistinct can be verified similarly as in Part (c). Therefore, \ud835\udf0e2 can again be selected as \ud835\udf0e2 = + 1 without loss of generality. (h) Position Singularities Associated with the Slider Position For particularly specified values of s03, there may occur two kinds of position singularities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002381_chicc.2016.7554554-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002381_chicc.2016.7554554-Figure1-1.png", + "caption": "Fig. 1: Simple-pendulum system driven by a DC motor", + "texts": [ + " In this paper, we first provide the mathematical model for multiple simple-pendulums driven by DC motors and present the consensus tracking problems for multiple simplependulum network systems. Then the distributed adaptive consensus tracking control protocol is designed and the stability condition is deduced. Simulation results are supplied and prove the effectiveness of proposed method, and finally the conclusions are drawn. 2 System modeling and problem formulaiton The structure of a simple-pendulum system driven by a DC motor is shown in Figure 1. Herein, u is the terminal voltage on the armature winding. R and L are resistance and inductance of the armature circuit. ia is the armature current and E is back electromotive force of the DC motor. The swing angle of simple-pendulum and the quality of the ball are \u03b8 and m, respectively, and l is the length of the pendulum rod. Each simple-pendulum driven by DC motor accordingly is taken for each agent, and then we call the multiple simplependulum network system as a multi-agent system, which can be expressed as a set\u03a9 = \u03c9i, i = 1, 2, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001795_robio.2015.7418892-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001795_robio.2015.7418892-Figure1-1.png", + "caption": "Figure 1. Kinematic modeling of 7 DOF manipulator.", + "texts": [ + " Section III describes the method of trajectory planning in Cartesian space. The trajectory planning theory applied to manipulator is tested on Section IV. In Section V the final outcome of this work is briefly discussed. In this paper, SHUNCK 7-DOF manipulator is used, which consists of 7 rotary joints and one translational joint. Every rotary joint has one DOF, and the adjacent axes are vertical arranged. The terminal gripper has one open-close translational DOF. The coordinate system of each joint is displayed in Fig 1. D-H Method [12] describes the attitude relationship of two adjacent coordinate systems with homogeneous transformation matrix, in which the four parameters are: the common normal length 1ia from 1iZ to iZ measured along iX , the included angle 1i from 1iZ to iZ measured about iX , the distance id from 1iX to iX measured along iZ , the rotational angle i from 1iX to iX measured about iZ . The D-H parameters of initial status are shown in Table 1. 978-1-4673-9675-2/15/$31.00 \u00a9 2015 IEEE 940 Expressed 1i iT as homogeneous transformation matrix that relates coordinate system { i } to coordinate system { 1i }" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003445_j.procir.2020.05.186-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003445_j.procir.2020.05.186-Figure4-1.png", + "caption": "Fig. 4. Results of the stripe light detection of the ground surface according to the CAD model", + "texts": [ + " The accuracy of the measurement report depends on the accuracy of the used five-axis milling machine. To avoid a possible error due to calibration, the geometric features of the test specimen were measured by an optical 3D Scanning system with Blue LED Fringe Projection (ZEISS Comet LED, Carl Zeiss AG, Oberkochen, Germany). The obtained results were compared to the tactile measurement report. It was determined, that the maximum deviation of the ground surface in this comparison is 0.0602 mm. The results of the optical measurement are shown in figure 4. It was demonstrated, that there is a deviation between optical and tactile measurement. The reasons might be: \u2022 The test specimen was made of aluminium. For the measurement it had to be coated with a titanium dioxide spray due to the highly reflective surface. \u2022 Although this coating was necessary, the stripe light pro- jector results are undersized. \u2022 These results could have been falsified by an incorrect calibration. In a second step the measuring method was applied to an AM turbine blade, that is made of Inconel 718 by laser beam melting (LBM)", + "000 6.018 0.018 4 6.000 6.015 0.015 5 6.000 6.015 0.015 To avoid a possible error due to calibration, the geometric features of the test specimen were measured by an optical 3D Scanning system with Blue LED Fringe Projection (ZEISS Comet LED, Carl Zeiss AG, Oberkochen, Germany). The obtained results were compared to the tactile measurement report. It was determined, that the maximum deviation of the ground surface in this comparison is 0.0602 mm. The results of the optical measurement are shown in figure 4. It was demonstrated, that there is a deviation between optical and tactile measurement. The reasons might be: \u2022 The test specimen was made of aluminium. For the measurement it had to be coated with a titanium dioxide spray due to the highly reflective surface. \u2022 Although this coating was necessary, the stripe light pro- jector results are undersized. \u2022 These results could have been falsified by an incorrect calibration. 0,00 - 1,00 1,00 0,20 0,40 0,60 0,80 - 0,20 - 0,40 - 0,60 - 0,80 Fig. 4. Results of the stripe light detection of the ground surface according to the CAD model In a second step the measuring method was applied to an AM turbine blade, that is made of Inconel 718 by laser beam melting (LBM). According to [5], the position of a part in the additive manufacturing machine is crucial for the geometric accuracy. Therefore, turbine blades were measured with the developed measuring strategy. Although all turbine blades were produced in one batch and on the same machine, there are differences in the geometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001049_ijat.2013.053164-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001049_ijat.2013.053164-Figure3-1.png", + "caption": "Figure 3 A manual gun drill resharpening apparatus consists of a three-axis fixture and a grinding spindle with fixed axis (see online version for colours)", + "texts": [ + " The rotational speeds and feed rates of the drills were kept consistent at 700\u2013900 rpm and 5\u20139 mm/min respectively. Coolant pressure was maintained at 6 MPa. After each drilling cycle of 50\u201390 mm, the cutting efficiency reduced gradually with the development of tool wear on the drills primarily on the bearing pads as well as the rake and flank faces of both the inner and cutting edges. The drills would be retrieved from the partially drilled hole, dismounted from the machine and then resharpened manually by the operator with a manual grinding apparatus as shown in Figure 3. Regeneration of the drill geometries relied on a three-axis gun drill fixture. The fixture was mounted to a universal tool grinder with a single-speed work spindle and fixed grinding axis. Cup-type, vitrified diamond grinding wheel was used for material removal with the intention to maintain the flatness of the cutting edges. The drill was mounted to the gun drill fixture with the secure of the carbide tip on a V-clamp and the support of the long drill stem with an extended balancing rest. The main geometries to be regenerated in sequence were the primary relief, the secondary relief, the inner relief, the oil clearance, the front clearance and the lead-in chamfer as shown in Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002946_j.precisioneng.2020.03.001-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002946_j.precisioneng.2020.03.001-Figure8-1.png", + "caption": "Fig. 8. Schematic of the two top specimens and the base specimen: (a) the two top specimens, (b) the base specimen.", + "texts": [ + " If the waviness amplitude w can be made smaller than \u03b5dc, all the real contact areas per crossing becomes larger than amin. Section 2 describes the conditions for distributing the real contact areas uniformly in the nominal contact area using the CMC method. In this section, we report that the distribution of the real contact surface can be controlled successfully using the CMC method and that this was verified by experiment. Three specimens were made for the contact experiment, two top specimens and a base (bottom) specimen (specimen 1, specimen 2, and the base). Fig. 8 shows a schematic of the top specimens and the base specimen. Specimen 1 had minimal waviness within the condition of uniform distribution of the real contact areas. Specimen 2 had a step and did not satisfy the condition of uniform distribution of the real contact areas. The base specimen had the same surface shape as specimen 1. Surface contact was made by pushing the top specimens onto the base specimen and then applying a preload. The distribution of the real contact areas on these contact surfaces was investigated by measuring the plastic deformation of the contact surfaces with a laser microscope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001020_icems.2014.7013720-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001020_icems.2014.7013720-Figure4-1.png", + "caption": "Fig. 4. The 3D finite element model of PMLSM.", + "texts": [ + " Through the above analysis, the d- and q-axis inductance of PMSLM can be obtained only by measuring the basic electrical parameters at two special rotor positions when arbitrary two phase windings are energized by alternating current. Fig. 3 shows the two special positions for measuring the d- and q-axis inductance of PMSLM. Coincidentally, the line voltage reaches its extreme value at the two positions. Therefore, we only need to detect the line voltage instead of the accurate positioning in advance. To obtain reference values of inductance, a 3-D finite element model of PMLSM, shown in Fig. 4, is built in commercial FE software JMAG. The parameters are shown in Tab. I. Fig. 5 shows the simulation results of self-inductance and mutual inductance and the values are shown in Tab. II. The inductance is basically constant when the electrical degree changes because of the non-salient pole feature. So, the second harmonic component is zero and Ld is equal to Lq in this case. In addition, the three phase inductances are not exactly equal due to the end effect of linear motor. According to (1), the dand q-axis inductance from finite element method is 0 0= =36" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001689_s11071-015-2356-y-Figure13-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001689_s11071-015-2356-y-Figure13-1.png", + "caption": "Fig. 13 Characteristic curves of the possible flowmeter. Left panel for \u03b7 < j , the arrows denote the hysteresis effect. Right panel \u03b7 > j , coincidence between curves of the different cases aremade unambiguous by the dotted lines. The black dot denotes the location of transient effects at unloading", + "texts": [ + " For example, all the bifurcations in Table 5 are codimension-1, and the codimension of the corresponding discontinuity sets are denoted in brackets. 6.2 Bifurcation diagrams In the different kinematic cases, the equilibrium solutions are located in different subspaces of X . Thus, for the full bifurcation diagram of\u03c90, a seven-dimensional graph would be needed. If we want to describe the location of the equilibria by a single variable, the natural choice of \u03c9c does not result in clear bifurcation diagrams due to overlapping of curves (see Fig. 13). Instead, let us define \u03c90 x := \u03c9\u03030 x + R r \u03c90 c , (91) where\u03c9c \u2261 \u03c90 c and \u03c9\u0303x \u2261 \u03c9\u03030 x are the coordinates of the stationary solution. Physically, (91) means the angular velocity of spinning on the wall [the first coordinate of (4)]. Table 5 Bifurcation values of \u03c90 with the name and abbreviation of the bifurcation and the corresponding kinematic cases Value Bifurcation Notation Cases \u03c901 Persistence (codimension-3) P3 RR\u2013SS \u03c902 Nonsmooth fold (codimension-2) NF2 SS\u2013SR \u03c902 Persistence (codimension-2) P2 NS\u2013NR \u03c903 Nonsmooth transcritical (codimension-1) NT1 RR\u2013SR \u03c904 Fold F SS Equilibria of the system can be visualised more clearly if they are divided into two groups", + "3 Limitations of operation of the flowmeter As it was mentioned in the introduction, operation of the flowmeter is based on measuring the velocity of the ball around the vessel, which means measurement of \u03c9c(t). For stationary flow velocity characterised by\u03c90, a constant \u03c9c(t) \u2261 \u03c90 c value is measured if we are in the bases of attraction of the stationary solutions. In the descriptions of many patents, it is explicitly assumed that the ball remains in the case of dual-point rolling (Case RR) with \u03c90 c (t) \u2261 \u03c90, but we have seen that the situation is much more complicated. In Fig. 13, we can see theoretical characteristic curves of the flowmeter which are based on the results of the previous subsections. In the left panel of Fig. 13, it can be seen that in the case \u03b7 < j , the dual-point slipping causes complications. At \u03c90 = \u03c901, the stationary solution passes over to Case SS, and there is a jump in the slope of the characteristic curve. Between \u03c901 and \u03c902, calibration of the flowmeter is possible, but for physically realistic parameters, the slope of the curve is very small here, which decreases the accuracy of the flowmeter. Between\u03c902 and\u03c904, two co-existing stable stationary solutions can be found. This causes a hysteresis effect (denoted by arrows in Fig. 13) and makes the calibration impossible in this region. In the case \u03c90 > \u03c902, the co-existing stationary solutions of Cases SR and NR correspond to \u03c90 c = \u03c90, which can be used for measurement, again. On the right panel of Fig. 13, the case \u03b7 > j can be seen. The stationary solutions of Cases RR, SR and NR all overlap on the line \u03c90 c = \u03c90, which results in a more favourable dynamics for the flowmeter. However, the motion corresponding to Case RR is different from the others (see also Fig. 11), because other coordinates of the solution do not coincide. If in the region \u03c90 > \u03c903, a small perturbation pushes the stationary solution from Case SR to Case NR, the dynamics can remain in this case also for slow decreasing of \u03c90 between \u03c902 and \u03c903. As the characteristic curves of Cases RR and NR coincide, this hysteresis effect is interesting only when we reach \u03c902 by decreasing \u03c90 (see the black dot in Fig. 13). Then, the stationary solution of Case NR vanishes, and after a temporary existing solution of Case NS (see Fig. 9), transient effects and impacts can occur before reaching the stationary solution of Case RR. This is, again, unfavourable from the point of view of the measurement. To summarise, we can say that in both cases of Fig. 13, transitions between different kinematic cases and bifurcations cause limitations in favourable operation of the flowmeter. To keep the stationary solution in Case RR, one possible strategy is to increase the values of\u03c901 (if\u03b7 < j) and\u03c902 (if\u03b7 > j) through theparameters of the system. Alternatively, shape of the swirling blades could be tuned to reach not too large values of \u03c90. Toomuch reduction in\u03c90, however, results in inaccuracy in the measurement of the velocity of the ball. Let us emphasise that our analysis was restricted to the stationary solutions in the case of stationary fluid flow" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000914_20130825-4-us-2038.00108-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000914_20130825-4-us-2038.00108-Figure11-1.png", + "caption": "Figure 11. The geometry and notations for the face during bucket filling.", + "texts": [], + "surrounding_texts": [ + "While the forces from the weight and inertial components of the machine can be accurately calculated with the presented model from catalog data, the digging force on the bucket is dependent on many factors. Hemami et. al (1994) present a model for this force based on a published formula. However, the force in their formula is a simple, linear function of the cutting edge angle, and the other constants need calibration. A new approach was followed, using force and mass balance analysis in the bucket and in the digging face. Figures 10 and 11 show the geometries and notations for the bucket and the face, respectively, during bucket filling. From fiction in the bucket and lifting of center of mass, the horizontal force is: (21) where (22) Substituting (22) to (21) gives the maximum value for force F: (23) From fiction in the muck pile along friction planes according to , the force balance equation is : (24) In equation (24), V is the volume of the material being sheared and pushed up: (25) The additional force in Eq. (25) is a horizontal force component to lift the amount of mass, expressed from energy balance: ( ) (26) where (27) (28) Substituting (25)-(28) into (24) gives: (29) The force F(x) in equation (23) and (29) are equal under slow, continuous motion, neglecting initial force, and under the condition of a positive F for the bucket. If this condition does not hold, the result will assume tensile force in the muckpile, a physical impossibility. This scenario must be checked before the balance equation is used for solution. After rearrangement, a quadratic equation is obtained for H(x): ( ) , where: (30) ( ) (31) (32) [ ] (33) The vertical distance of the center of mass in the bucket ( is measured from the bucket digging edge. The value of depends on the angle of the digging direction , the bucket angle , and the shape and the instantaneous fill factor, Cm, of the bucket. An approximate formula can be derived for simplified bucket geometry as follows: (34) In Eq. (34), is instantaneous mass of loaded material, M is the mass capacity of the bucket, K is the distance of the line of the center of gravity along which the material moves within the bucket during loading, and L is the opening height of the bucket. The ratio in Eq. (34) m/M is the instantaneous fill factor, Cm: (35) The solution for Eq. (30) provides the height of the material, H, right in front of the bucket loading edge: \u221a (36) Height, H, is a function of several parameters in Eqs. (31)- (34). Two groups of parameters can be defined: Group P1, involving only constant material properties, and group P2, involving digging control and face geometry parameters which are all variable: (37) ( ) (38) The operator of the hybrid machine in \u201cbucket steering\u201d kinematics directly controls and , while is determined by the shape of the fallen face, all variable with the instantaneous positions of the bucket edge. The question is how the bucket filling process plays out from the spotting point at Cm=0, to its completion to Cm=1. In other words, how will Cm(x) depend on (x), (x), and (x)? It must be pointed out that the ultimate hybrid control would be the one which directly controls Cm from 0 to 1 with ease and minimum energy, time and machine wear. In order to connect H to the movement of the bucket along x, another equation is needed. It must be realized that the material slides into the bucket and it does not directly follow its displacement. It is possible to define two elastic parameters, one for the material in the bucket, kB, and one for the pile, kF, as the derivatives of the horizontal forces with respect to x. From Eqs. (23), (29) and (33) we obtain: ( ) (39) (40) The bucket is moved by dx which causes material movements, dxB into the bucket and dxF into the face: (41) With the kB and kF elastic constants, dxB can be expressed: (42) Substitution of Eq. (42) into Eq. (41) and simplifications give: (43) Since , dm can be expressed with the use of Eq. (43): (44) Substituting ycp from Eq. (34) yields, after rearrangement, a differential equation for m as a function of x: (45) Note that m/M=Cm in Eq. (45). The solution for the digging force by moving the bucket is now complete. Starting with a small m and H, first, the dm mass has to be determined for H from the solution of Eq. (45). Next, the new H value can be evaluated from Eq. (36), and the process can continue along the digging trajectory. The digging force F(x), is given by Eq. (29), assuming a positive value for the bucket. If this condition does not hold, the result predicts tensile force in the muckpile, a physical impossibility. This scenario must be checked and corrected by setting the digging depth, H, as a free parameter under direct motion control before Eq. (29) is used for the solution. The procedure was coded into the Hitachi 3500 excavator analysis model in Matlab. The analysis results are shown in Figs. 12 trough 14. The model simulated the motion of the machine along a given trajectory of the bucket edge and variable bucket angle, shown in Fig. 12. Figure 13 shows the joint torques from the lowest to the highest point along the digging part of the trajectory. Figure 14 depicts the cumulative work generated by the joint cylinders assuming 100% energy efficiency and no recuperative energy recovery during lowering of any joints during bucket loading. In the sections of clockwise (lowering) motion, the energy was equated with zero." + ] + }, + { + "image_filename": "designv11_34_0001644_ijrapidm.2015.073548-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001644_ijrapidm.2015.073548-Figure1-1.png", + "caption": "Figure 1 Geometry of the case study: top, side, front and perspective view", + "texts": [ + " The geometry includes a variety of macro-level geometrical features machinable in a three-axis milling process, such as curved surfaces, holes, pockets, sharp edges, as well as perpendicular, parallel and inclined planes at different angles. The geometry was designed to minimise the need for support structures by eliminating possible overhanging features, and the functional geometry was located only in the XY plane in order to simplify the milling process, which was completed with a three-axis milling machine with the same tool in a single set-up. Figure 1 shows the geometry used in this experiment. The part in grey shows the nominal geometry. Overlapped in the same figure, a translucent offset geometry is displayed. The nominal geometry has been offset in all faces and geometrical features to compensate for the overall distortions of the additive process. By doing so, the parts produced will have enough stock material to be removed during the post-processing; therefore, inaccuracies of the AM parts can be compensated for during the finishing process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000826_s1052618814050136-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000826_s1052618814050136-Figure4-1.png", + "caption": "Fig. 4.", + "texts": [ + " Thus, altogether we get 15 elements and 16 one degree of freedom kinematic pairs in the mechanism. Substituting values in (1) we shall get W = 6 \u00d7 (15 \u2013 1) \u2013 5 \u00d7 16 = 4. This result is true: the number of degrees of freedom is four. Let us consider the mechanism, being the evolution of the \u201cORTHOGLIDE\u201d robot, with two added rotary degrees of freedom; and rotation is transmitted to the same kinematic chains that provide transla tional motion. We use formula (1) for spatial mechanisms to determine the number of degrees of freedom of the mech anism with five degrees of freedom (Fig. 4). Let us calculate the number of elements n in the mechanism and number of the one degree of free dom kinematic pairs p5. The spatial mechanism (Fig. 4) includes base 1, output element 2, working organ 3 and three kinematic chains. Each chain contains a sliding motor 4, 4 ', 4 ''; an initial rotary kinematic pair 5, 5 ', 5 ''; initial element of the parallel crank mechanism 6, 6 ', 6 ''; rotary kinematic pairs of the par allel crank mechanism 7, 7 ', 7 ''; terminal rotary kinematic pair 8, 8 ', 8 ''; the terminal element of the par allel crank mechanism 9, 9 ', 9 ''; intermediate elements of the parallelogram and terminal element of the kinematic pair 10, 10 ', 10 ''", + " In this case, we take into account the same elements that were used before, excluding intermediate ele ments of parallelogram. Three progressive kinematic pairs will change twelve sliding kinematic pairs of three parallelograms. According to (1) we shall get W = 6 \u00d7 (15 \u2013 1) \u2013 5 \u00d7 16 = 4. This result is not true as well. 382 JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 43 No. 5 2014 NOSOVA et al. Let us change the parallel crank mechanisms in kinematic chain not containing a rotary motor (Fig. 4, III) by a universal joint (Fig. 4, IIIa) and we shall consider the parallel crank mechanism as the sliding kinematic pair in the other two kinematic chains containing rotary motors. Thus, altogether we get 16 ele ments. The number of 5 one degree of freedom kinematic pairs p5 is 17. W = 6 \u00d7 (16 \u2013 1) \u2013 5 \u00d7 17 = 5. This result is true: the number of degrees of freedom is five. Let us consider a mechanism with six degrees of freedom and we use formula (1) for spatial mecha nisms to determine the number of degrees of freedom of the manipulating mechanism (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002990_012029-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002990_012029-Figure2-1.png", + "caption": "Figure 2. Normal vector to a teeth surface at contact point.", + "texts": [ + "A A B BJ J ( ) ( ) \u00d7 - = \u00d7 - = - E r r E r r E E e e e B A B e A A A A r f 0 3 1 0 0 3 E0 E e eA A 3 3 eA3 yA eA1 eAref1 ToPME-2019 IOP Conf. Series: Materials Science and Engineering 747 (2020) 012029 IOP Publishing doi:10.1088/1757-899X/747/1/012029 The value of angle and the radius vector are determined as Basically, the vectors and indicate at different contact points. The difference between contact points determine the vector of elastic mismatch Also a normal vector to a teeth surface at contact point is defined as , , the pressure, cone and helix angels respectively (Fig.2). Scalar product of and determine the normal contact offset . y = \u00d7 y = \u00d7 e e e e cos sin A ref A r A A A efA 1 1 1 2 yB eBref1 r y = y r = y + ye e ecos sin A B A B B B B ref B B1 1 2 eAref1 eBref1 ( ) ( )D = + + r - + + ru r u e r u eB A B B B ref A A A ref0 1 0 1 + += = g \u00d7 a - a \u00d7 b \u00d7 g = a \u00d7 b = a \u00d7 g \u00d7 b + a \u00d7 g n e e e cos sin cos sin sin cos cos cos cos sin sin sin A A A ref reff f f f f f 1 1 2 2 3 3 1 2 3 a b g Du n D = \u00d7 Dn us ToPME-2019 IOP Conf. Series: Materials Science and Engineering 747 (2020) 012029 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001925_ceidp.2014.6995760-FigureI-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001925_ceidp.2014.6995760-FigureI-1.png", + "caption": "Fig. I. 3 D model of composite tower", + "texts": [], + "surrounding_texts": [ + "been widely used in transmission lines of different voltage ratings. Composite tower has advantage in saving line corridors, reducing tower weight, easy installation and transportation when compared with all-steel towers. It has got a growing attention from researchers. Owing to its complex structure, satisfactory operation of composite tower is intimately related to the potential and surface electric field distribution. With a three-dimension (3- D) finite-element method (FEM) software ANSYS, this paper presents computation results of the potential and electric field distribution for composite tower in 330 kV double circuit transmission line. The influence of tower, conductors, grading rings and ground has been analyzed. The voltage undertaken by cross-arms and tower body has been calculated in different situations. A set of grading rings configuration program has been proposed to control the electric field strength of the position that need to be concerned. After optimization, the electric fields strength of the key location of composite tower has been proved that they could meet the critical values.\nKeywords-composite tower; electric field distribution; finite\nelement method(FEM); grading ring\nI. INTRODUCTION\nWith the purpose of increasing power transferred and improvement of the efficiency of power delivery, the system voltage rating has been increasing steadily [1]. In northwest of China, the 330 kV grid is very important, it is related to the safe and stable operation of northwest grid and new energy development.\nCurrently, the steel transmission tower is widely used in power transmission lines. The great demand for steel consumes a lot of mineral resources in the construction of transmission lines and the steel production and processing harm eco-environment hugely. Meanwhile, the steel has shortcomings in heavy weight, easy to rust, poor low temperature performance and so on.\nDue to its excellent anti-pollution performance, light weight and ease of installation [2], composite materials have been widely used in various transmission lines of different voltage levels [3-4]. Among of them, composite tower is a smart application. Usually, insulator strings are used to connect tower and wires to provide electrical insulation and mechanical support. If the steel cross-arms and the middle part\nof the tower is replaced with composite materials, the new type composite cross-arms can also take the role of insulators, thus the insulator string can be cancelled. Compared with the steel transmission tower, this new composite tower has advantage of saving line corridors, reducing tower weight, easy installation and transportation.\nComposite cross-arms and towers have already been used in Europe and America. Several scholars have done research on them [5-7]. However, the applications and researches above mainly focus on the mechanical behavior rather than electrical performance. Furthermore, the tower structure is relatively simple. In contrast, the 330 kV double circuit composite tower has more complex structure, more uneven potential and electric field distribution.\nThe role of grading ring is to uniform the potential distribution of composite cross-arm, decrease the electric field strength on the surface of sheath and prevent the corona discharge on the surface of fittings. For composite insulators, the electric field magnitude at 0.5 mm above the surface of the sheath should be less than 450 V/mm (rms) and a surface electric field 2500 V/mm is often used as a reference value for design purposes [8]. This paper adopts the critical electric field values based on the reference above to optimize the design of composite tower.\nUsing the FEM software ANSYS, three-dimensional (3-D) FEM models of 330 kV double circuit composite tower based on actual sizes is presented in this paper. The aim of the paper is to analyze the structure of 330 kV double circuit composite tower thorough the computation of potential and electric field distribution and then apply to the real project based on the findings herein.\nThe tower, with a height of 100 m, is a double circuit line tower. The base and head parts of the tower are made of steel, the middle part of the tower and the cross-arms are made of composite materials. As shown in fig. 1, the three phases are\n978-1-4799-7525-9/14/$31.00 @20141EEE 283", + "named phase A, phase B and phase C, respectively. The length of phase A is 10 m with an installation height of 38 m; the length of phase B is 11 m with an installation height of 33 m; the length of phase C is 10 m with an installation height of 28 m. The middle part of the tower is mainly made up of four vertically arranged post insulation rods. These post insulation rods are divided into four sections by three intermediate flanges. Adjacent post insulation rods are connected by oblique insulation rods.\nThe structure of composite cross-arms in three phases is similar. As shown in fig. 2, the main part of the composite cross-arms is a two V -shaped post insulators, and two upward sloping V -shaped tension insulators, Post insulators and tension insulators are connected by connecting fittings and connected to the tower by connecting fittings. Post insulators are connected with conductors by yoke plate. The conductor used here is four bundle, the subconductor diameter is 26.8mm and the bundle diameter is 400 mm.\nConsidering the interaction between three phases, the harmonic analysis method was used to obtain the maximum electric field strength in one entire power frequency cycle. In\ncalculation, the phase angle of phase B is assumed to be on 0\", the potentials of excitation sources are prescribed considering the maximum instantaneous values, so 296.4 kV is applied on\nphase B, -148.2\u00b1i256.7 kV are applied on phase A and C, respectively. The electric potential is set to zero for the base and head parts of the tower. The three intermediate flanges can be seen as floating potential determined by the whole potential distribution.\nTable 1 shows the maximum voltage undertaken by composite insulators, post insulation rods during one entire power frequency cycle. It can be seen that post insulators undertake about 73.3 % to 87.1 % of total voltage; considering the interaction between three phases, tension insulators undertake about 100% to 108.7 % of total voltage; post insulation rods undertake about 27.1 % to 33.3 % of total voltage. Considering the situation that the middle part of tower was grounded, than the potential of intermediate flanges is equal to zero. At this time, composite cross-arms undertake the whole voltage, the voltage undertaken by post insulators\n978-1-4799-7525-9/14/$31.00 @20141EEE 284", + "increase about 12.9 % to 26.7 % of total voltage, while the voltage undertaken by tension insulators decrease about 0 % to 8.7 % of total voltage.\nThe potential distribution curves taken thorough the center of post insulator and tension insulator in phase A are shown in fig. 5. The potential distribution inside post and tension insulators is basically the same. The curve near flange drops significantly when compared to the middle part, it indicates the high voltage drop near flange.\nFig. 6 shows the electric field distribution on the surface of composite cross-arm in phase A. Similar results can be derived from the other two phases. It can be seen that the place near the end of the silicone rubber sheath suffers much higher electric field strength than the middle. Table 2 gives the maximum electric field strength on the surface of sheath and fittings in three phases. It can be seen that the maximum electric field strength on the surface of end fittings and yoke plates is below 2500 Vlmm, it satisfies the critical value. The maximum electric field strength in the line end of post insulator is about 600 Vlmm, the maximum electric field strength in the line end of tension insulator is about 2300 Vlmm, which exceed the critical value, the sheath is likely to be damaged due to electric ablation during long time running. Consequently, initial design of composite cross-arms needs to be optimized.\nThe installation of grading rings needs to ensure the maximum electric field strength on the surface of the end fittings and silicone rubber sheath below the critical values. Besides, the grading rings themselves do not produce corona and are easy to install. After a number of calculation and optimizations, the final installation program is presented herein. As shown in fig. 7, grading rings with different sizes are installed at the line end of post, tension insulators and the tower end of tension insulators to shield the electric field on the surface of sheath.\nThe calculation results after optimization are illustrated in table 3, 4 and fig. 8. As shown in table 3, the whole potential distribution of composite tower almost unchanged after optimization, while the potential distribution near the line end of insulators improves. As shown in fig.8, after the installation of grading rings, the region of high electric field transfers from the surface of sheath to the surface of grading rings and the maximum surface electric field strength of sheath decreases\n978-1-4799-7525-9/14/$31.00 @20141EEE 285" + ] + }, + { + "image_filename": "designv11_34_0001107_icems.2014.7014032-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001107_icems.2014.7014032-Figure9-1.png", + "caption": "Fig. 9 Coil arrangement of the wound type rotor.", + "texts": [ + " To decide the coil configuration of the wound rotor on the basis of the magnetic flux density distribution due to the magnet, magnetic flux density distribution by magnet was obtained by replace the original stator to the slot-less stator which shown in Fig. 3. The number of phases is six by considering the slot harmonics. The coil pitch decided as 30 degree (electric angle). The winding uses concentrated winding. The distance of each coil from the center of the rotor is determined based on magnetic flux density distribution due to the magnet. The distance from the center of the rotor of each coil is determined by changing the distance every one millimeter. Fig. 9 shows the coil arrangement of the wound type rotor. Fig. 10 shows the original and reproduced magnetic flux density distribution of the air gap created by the magnet or wound type rotor. As shown in Fig. 10, it is possible to reproduce the magnetic flux density distribution due to the magnets by wound type rotor. However, the problem of harmonics has remained. Fig. 11 shows the magnet motive force due to current which is inputted to each coil. As shown in Fig. 11, large magnet motive force is required to reproduce the magnetic flux density distribution due to the magnets since the rotor uses non-magnetic material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003282_s10015-020-00616-4-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003282_s10015-020-00616-4-Figure3-1.png", + "caption": "Fig. 3 Experimental setup for motion performance testing", + "texts": [ + " We conducted two experiments in a swimming pool to test the motion performance of the towed ROV. One experiment investigated whether the vehicle depth can be adjusted by changing the main wing angle while maintaining the tail wing angle at 0 degrees. Another experiment was undertaken to investigate whether the vehicle pitch can be adjusted by the angle change of the horizontal tail wings with the main wings fixed at 0 degrees. The tail wings serve in an important role for adjusting the angle of attack of the main wings. The experimental setup is presented in Fig.\u00a03. The initial main and tail wing angles were 0 (\u00b0). Both wing angles have a software minimum and maximum limit of \u00b1 20 degrees. The towed ROV contacted the bottom of the swimming pool in the initial state (t = 0). Instead of towing the vehicle by boat, a person pulled the vehicle with a rope during the experiment. Figure\u00a04 depicts time series data of the depth and main wings of the towed ROV when the angle of the main wings was changed by the operator\u2019s manual operation. The depth was calculated from the measured values of the laser distance sensor used in an earlier study [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000581_s40430-013-0118-7-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000581_s40430-013-0118-7-Figure1-1.png", + "caption": "Fig. 1 a Three double row four-point-contact ball bearings mounted on a wind turbine as pitch bearings. b Elements of a rotational connection with a single row four-point-contact ball bearing", + "texts": [ + " The relationship between groove shape and clearance variation is established in the paper and two experiments are designed to verify the correctness of theoretic analysis. A new processing technology for final processing of groove in four-point-contact slewing bearings is suggested according to theoretic analysis and experiment research in the paper. Keywords Slewing bearing Four-point contact Groove shape Starting friction torque Negative clearance Slewing is defined as the rotation of an object around an axis [1\u20133]. Thus, a slewing bearing (as shown in Fig. 1a) is a bearing used in slewing applications for transferring axial force Fa, radial force Fr, and turn over moments M singularly or in combination [2\u20138], as can be seen in Fig. 1b. The inner and outer rings have mounting holes that allow the bearing to be bolted directly to the supporting structures [9]. The balls or rollers are inserted into the bearing through a radial cylindrical hole in one of the rings [3]. The hole then is closed using a removable loading plug contoured to the ball path or roller path surface. The most common materials for rings of slewing bearing is GB-42CrMo4 alloy steel (equivalent to DIN 42CrMo4V) and the most common materials for balls or rollers is GBGCr15 (equivalent to DIN 100Cr6) [10]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001696_978-3-319-20463-5_22-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001696_978-3-319-20463-5_22-Figure3-1.png", + "caption": "Fig. 3 The test bed (a), location of pinions (b) and sensors (c)", + "texts": [ + " The use of the Fourier transform on an angularly sampled rather than time-sampled signal overcomes the rotating speed variations and allows to consider directly the characteristic frequencies of the different signals. For instance, the frequencies of observation of the gears are no more changed by the rotational speed of the machine, but are directly observable peaks corresponding to the number of gear teeth [5]. Thus, the frequency channel of interest for the gear will be directly identified from the kinematics of the machine and the acquisition parameters (resolution of optical encoder and length of acquisition). The test bed Fig. 3 used in this study consists of two rotating shafts, on which are mounted a pinion and a spur gear offering a gear ratio of 25/56. To compare the effectiveness of methods of analysis, we used six pinions with different fault states. The first one is referred as Good (G), whereas the others have several different types of defects: a Root Crack (RC), a Chipped Tooth in Width (CTW), a Chipped Tooth in Length (CTL), a Missing Tooth (MT), and General Surface Wear (GSW) as shown in Fig. 2. Three pinions are simultaneously mounted on the input shaft of the gearbox, the engagement change is done by a simple axial movement of the wheel on its axis Fig. 3b. The input shaft is driven by an electric DC motor controlled in rotational speed. The engine ensures a maximum speed of 3600 rpm. The output shaft is connected to a magnetic powder brake capable of generating different resistive torques. To record vibration signals, two accelerometers (sensitivity: 100 mV/g) are mounted radially, one vertically and the other horizontally on the outer surface of the bearing case of the output shaft of the gearbox as shown in Fig. 3c. To measure the angular positions of the shafts, two optical encoders of 2500 pulses per revolution are mounted at the free ends of the two shafts of the gearbox Fig. 3a. The clock frequency of the counting acquisition system is 80 MHz, generally considered sufficient to locate the rising edges of the encoder signals. The time sampling frequency of the accelerometer channels is 125 kHz. The cutoff frequency of the anti-aliasing filter is 27 kHz. The acquisition duration is 30 s. The accelerometer signals and the angular positions have been recorded for different operating conditions by varying the rotation speed and the resistant torque for each of the six gears used (Table 1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003051_s10409-020-00937-4-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003051_s10409-020-00937-4-Figure1-1.png", + "caption": "Fig. 1 Elastica geometry and description. a Material point C is quasistatically controlled downward. b Whole elastica is regarded as an assembly of two pinned-clamped elasticas", + "texts": [ + " For shootingmethod, a series of numerical integrations in the iteration process should be applied repeatedly for all points on one stiffness\u2212curvature curve. To obtain the stiffness\u2212curvature curves efficiently, we firstly record typical pinned-clamped elasticas based on the second-order mode and then find pinned-clamped elasticas that correspond to different points on stiffness\u2212curvature curves quickly. The elastica is initially straight and buckles to the first-order mode under external forces at two ends, which induces the end-shortening. After that, two ends are pinned at A1 and A2 as shown in Fig. 1a. The material point C is quasi-statically controlled downward under force Q. Here, we regard the whole elastica as two components A1C and A2C as shown in Fig. 1b. Elasticas A1C and A2C can be regarded as two pinned-clamped elasticas and a general pinned-clamped elastica is shown in Fig. 2a. Then we study the shapes of the whole elastica by the stiffness\u2212curvature curves of two pinned-clamped elasticas, which describe the curvature at the clamped end as slope \u03b2R changes. To obtain the stiffness\u2212curvature curve of a pinned-clamped elastica, shapes of the elastica with different \u03b2R at the clamped end should be determined. Here elastica is inextensible and its weight is neglected", + " Horizonal location and deflection of the controlled material point are described by x0 and h. \u03b1 is the angle between A1C and A2C . The horizontal and vertical coordinates are x and y. \u03b8 is angle from horizontal to tangential direction. l is arc length of elastica A1A2 and l0 is the distance between A1 and A2. L is arc length of elastica AC and L0 is the distance between A and C . Subscripts 1 and 2 represent the quantities of elasticas A1C and A2C , respectively. The positive directions of the internal tangential force ft , normal force fn, and bending moment M are shown in Fig. 1b. The governing equations about geometrical quantities are \u03b8 \u2032 = k, x \u2032 = cos \u03b8, y\u2032 = sin \u03b8, (1) where k is the curvature. Prime represents differential with respect to the arc length s. Balance equations for the components of the internal forces and moment expressed in the Frenet\u2212Serret frame are ft \u2032 = k fn, fn \u2032 = \u2212k ft, M \u2032 = \u2212 fn. (2) According to themoment\u2212curvature relationshipM = E Ik, where E I is the constant bending stiffness, two quantities are defined: Ft = 2 ft/(E I ), Fn = 2 fn/(E I )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003405_rpj-07-2019-0200-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003405_rpj-07-2019-0200-Figure2-1.png", + "caption": "Figure 2 FFF testing geometry with dimensions and print strategy", + "texts": [ + " However, the porosity introduced by this reduced infill can be distinguished from the unintentional process-related porosity as shown in Figure 3. The LBPF samples were printed on an ORLAS Creator from O.R. Laser Technologie GmbH. The settings were 212.5W laser power, a track width of 57.5 mm, a beam diameter of 76.7 mm, a layer height of 30 mm and a scan speed of 1,200mm s 1. Both samples were printed using the aforementioned coreshell strategy with the part shapes scanned or extruded with two perimeter tracks and the core filled by a unidirectional hatching strategy. As can be seen in Figure 1(b) resp. Figure 2(b) the slicers used for each part utilized a slightly different strategy to generate the hatching. The LPBF part was sliced using the ORLas Suite Version 5.5.0.0 (O.R. Lasertechnologie GmbH, 2018) and generated distinct single scan paths for the exposure of the core of the part, whereas the FFF part was sliced using Slic3r 1.3.0 (Ranellucci and Lenox, 2018), which generated a single, connected path to fill up the core of the part. The LPBF print path contains considerably more single paths for this reason but also a longer total path length per area because of the significantly lower hatching distance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001204_aim.2013.6584161-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001204_aim.2013.6584161-Figure3-1.png", + "caption": "Fig. 3: Kinematically redundant 3(P)RRR mechanism with a maxmimum number of three additional prismatic actuators", + "texts": [ + " Therefore, the calculation time as well as the number of function evaluations are compared with respect to the proposed improvements and for different numbers \u03b7r of additional prismatic actuators, i. e. different dimensions of the optimization problem. Exemplarily, the motion planning of a kinematically redundant 3(P)RRR mechanism, being briefly reviewed in Sec. IV-A, is considered. It is important to note that the capability of kinematic redundancy, i. e. the potential of additional prismatic actuators, as well as the PSO algorithm, are not addressed in this paper. Besides others, the planar, kinematically redundant 3(P)RRR3 mechanism (see Fig. 3) is introduced in [7] and [27]. Basically, it is similar to the non-redundant 3RRR mechanism studied amongst others in [28]. Three kinematic chains GiMiPi (i = 1, 2, 3) connect the moving platform P1P2P3 to the base G1G2G3. In order to achieve kinematic redundancy, at least one prismatic actuator is added to the structure. This allows to linearly move the position rGi = (xGi , yGi) T of at least one active base joint Gi with respect to the actuator position \u03b4i. Depending on the degree of redundancy \u03b7r, the mechanism is driven by 3 + \u03b7r actuators" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003943_s00202-020-01146-9-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003943_s00202-020-01146-9-Figure1-1.png", + "caption": "Fig. 1 The manner of modeling different parts: a stator and rotor, b air-gap region", + "texts": [ + " In comparison with other MEC models developed for the SRM, the following merits can be highlighted for the MEC model introduced in the present paper: (1) The simplicity of the model because a low number of reluctances are used in the modeling and the airgap area is modeled by a continuous function dependent on rotor position, (2) accurate modeling of the saturation phenomenon especially in highly saturable parts of stator and rotor pole corners, (3) high computation speed useful for optimal design of the SRM where design stages need to be repeated frequently, (4) easy implementation of the model for various designs of different conventional types of the SRM such as 6/4, 8/6 and 12/8. In the following, the suggested model is described in section II. In order to validate this model, simulation and experimental results for a typical 8/6 SRM are presented in section III. Finally, the paper is concluded in section IV. In order to predict the phase flux linkage of the SRM, it is enough to consider the excitation of one phase and this is followed in the suggested MEC model as observed for an 8/6 SRM in Fig. 1a. Using three reluctances shown in Fig. 1b, change of air-gap reluctance due to rotation of rotor is considered in this modeling. The values of these reluctances are calculated using (1) to (3). Rg lg \u03bc0 Ag (1) Rgl lgl \u03bc0 Agl (2) Rgr lgr \u03bc0 Agr (3) where \u03bc0 is magnetic permeability of free space, Rg, Rgl and Rgr are middle, left and right air-gap reluctances, lg, lgl and lgr are average lengths of flux path in middle, left and right air-gap, and Ag, Agl and Agr are the cross-sectional area of flux path in middle, left and right air-gap, respectively", + " Based on estimation of the average cross-sectional area of flux path on stator and rotor poles, (10) and (12) are also obtained. With regard to Fig. 4, (10) to (13) can be justified. A \u2032 g wrp + hrp 8 l (10) l \u2032g \u221a r2si + r2ro \u2212 2rsi rrocos ( \u03b8 \u2212 \u03b2s + \u03b2r 2 ) (11) A\u2032 gr 0.5hrpl (12) l \u2032gr \u221a r2si + (rro \u2212 0.5hrp)2 \u2212 2rsi ( rro \u2212 0.5hrp ) cos(\u03b8 \u2212 \u03b1) (13) wherewrp is rotor pole width and the parameter \u03b1 is obtained as follows: \u03b1 tan\u22121 x rro \u2212 0.5hrp tan\u22121 rro sin \u03b2r 2 rro \u2212 0.5hrp (14) As observed in Fig. 1a, stator yoke and rotor core are modeled by Rsy and Rry calculated as follows: Rsy lsy \u03bc0\u03bcr Asy (15) Rry lr y \u03bc0\u03bcr Ary (16) where \u03bcr is relative magnetic permeability, Asy and Ary are cross-sectional area of stator yoke and rotor core, lsy and lry are the lengths of flux paths in stator yoke and rotor core, respectively. These parameters are obtained using the following equations. Asy hsyl (17) lsy \u03c0 ( rso \u2212 0.5hsy ) (18) Ary hryl (19) lr y \u03c0 ( rry \u2212 0.5hry ) (20) where hsy is the stator yoke height, rso is the lamination outer radius, hry is the rotor core height, and rry is the rotor core outer radius. In order to consider appropriately the local saturation, poles tips are modeled with two parallel reluctances as done in [18]. With regard to Fig. 1a, three reluctances (Rspm, Rsp1, and Rsp2) considered for stator pole can be calculated using the given equations: Rspm (2/3) \u00d7 hsp \u03bc0\u03bcrwspl (21) Rsp1 (1/3) \u00d7 hsp \u03bc0\u03bcr Asp1 (22) Rsp2 (1/3) \u00d7 hsp \u03bc0\u03bcr Asp2 (23) where wsp is stator pole width and Asp1, Asp2 are the crosssectional areas of flux at stator pole tip (overlapped and non-overlapped portions, respectively). Based on FE calculations, the stator pole flux widths in these two portions (wspf1 and wspf2 in Fig. 3a) are obtained for different rotor positions, and their variations are illustrated by the curves given in Fig. 3b. Using MATLAB Basic Fitting tool, the approximate functions to fit these curves are then determined and given in (24) and (25). Asp1 wspl(0.02\u03b8 + 0.016) (24) Asp2 wspl(\u22120.04\u03b8 + 1) (25) As done for modeling stator pole, three reluctances (Rrpm, Rrp1 and Rrp2) are also considered for rotor pole which are illustrated from Fig. 1a. Similarly, these reluctances are estimated using the given equations: Rrpm (2/3) \u00d7 hrp \u03bc0\u03bcrwrpl (26) Rrp1 (1/3) \u00d7 hrp \u03bc0\u03bcr Arp1 (27) Rrp2 (1/3) \u00d7 hrp \u03bc0\u03bcr Arp2 (28) Fig. 6 The parameters required to model the air-gap between the excited stator pole and the second rotor pole where Arp1, Arp2 are the cross-sectional areas of flux at rotor pole tip (overlapped and non-overlapped portions, respectively). Similar to that done above for stator pole, these cross-sectional areas are obtained for different rotor positions using the FE calculations, and the following approximate functions are then determined to fit the related curves: Arp1 wrpl(\u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002397_icma.2016.7558663-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002397_icma.2016.7558663-Figure3-1.png", + "caption": "Fig. 3 Anchor line model", + "texts": [ + " Turret system is directly connected with the FPSO vessel, and it is balanced through external forces and mooring system forces, the external forces can be achieved by thrusters through the controller. Each line can be divided into several elements, i the catenary theory is applied to simulate potential changes of the mooring system in the station-keeping process. The catenary calculation formulas is used at a fixed endpoint single mooring lines to solve the gravitational potential energy and the level of tension [11], and the anchor lines model is shown in Fig 3: For the lines of stress analysis by the catenary theory, the equation is: ( )11 2 / cosh 1 /s X h a h a h a\u2212\u2212 = + \u2212 + (3) where s is the length of the mooring line; X is mooring line horizontal projected length; h is the height of mooring line; a can be calculated as: /ha T w= (4) where, hT is the mass per unit length of the cable in the water, the unit is /t m . Specific calculation methods can be found in [12].In this paper, the total length of the anchor line is 2652 m and it can undertake the maximum tension 7553 kN " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002272_2016-01-1958-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002272_2016-01-1958-Figure6-1.png", + "caption": "Figure 6. Boundary conditions and loadings", + "texts": [ + " and knuckle The outer ring was modeled using SAE1055 material properties. Different material properties were applied to the heat treated area in comparison to the non-heat treated area. Elasto-plastic analysis was performed to compute the distortion behavior of the outer ring, bolts and knuckle. Figure 5 shows stress-strain curves which were obtained from uniaxial tensile testing. To consider the material property effect, a distortion analysis was performed with aluminum and a cast iron knuckles. As shown in Figure 6(a), fixed boundary conditions were applied to the bolt holes of the knuckle. The three degrees-of-freedom were fixed. To improve the reliability of the analysis, contact boundary conditions for the outer ring shield and knuckle were applied. Bolt-induced clamping torque, 180 Nm, was applied to the all three bolts shown in Figure 6(b)). A distortion analysis was performed by varying the concavity of the outer ring flange from 0 \u03bcm to 50 \u03bcm in 10 \u03bcm increments. Figure 7 shows the distortion contour at 0 \u03bcm. Maximum distortion occurs at the shield perimeter due to the size being both thin and large. It is important within this figure to focus on the outer ring distortion as key areas, that is, mounting location and raceways. Four points, outboard seal fitting point, P1, outboard raceway ball contact point, P2, inboard raceway ball contact point, P3, and inboard seal fitting point, P4, were selected from the outer ring because raceway contact points and seal mounting points are important for bearing performance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003367_s11432-019-2671-y-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003367_s11432-019-2671-y-Figure1-1.png", + "caption": "Figure 1 (Color online) Mechanical structure of the gliding robotic dolphin, showing the (a) conceptual design and (b) prototype.", + "texts": [ + " Our results suggest that we can free the flippers from having to control the pitch moment to focus on yaw control, potentially enabling us to decouple the control of three-dimensional (3D) motion. The remainder of this paper is structured as follows. Section 2 describes the gliding robotic dolphin\u2019s overall mechanical design and prototype. Next, we derive a full-state dynamic model in Section 3. Then, Section 4 discusses the control systems, including the LOS, CPG, and ACA methods. Section 5 discusses and analyzes the results of our simulations and aquatic experiments. Finally, Section 6 summarizes our conclusion and plans for future work. As illustrated in Figure 1, the gliding robotic dolphin has a streamlined shape, modelled after a killer whale, to reduce water resistance and obtain a better lift-to-drag ratio. Unlike previous gliding robotic dolphins [15,17], ours utilizes a more compact design, and replaces the external oil bladder with a water injector to make net buoyancy adjustments faster. In addition, the robot\u2019s shell is constructed from polyethylene. Table 1 gives details of our robotic dolphin\u2019s mechanical and electrical parameters. It has the following three main compartments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002508_j.ifacol.2016.10.495-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002508_j.ifacol.2016.10.495-Figure1-1.png", + "caption": "Fig. 1. Motion coordinate system", + "texts": [ + " To describe these motions, the following transfer functions are used: Wsway(s) = Kdv (\u03c4vs+ 1) , Wyaw(s) = 1 (\u03c4rs+ 1)s , (1) Wroll(s) = \u03c92 n s2 + 2\u03b6n\u03c9ns+ \u03c92 n , where Kdv, \u03c4v, \u03c4r, \u03b6n, \u03c9n are the fixed parameters. Taking into account the rudder forces and the disturbances, the vessel\u2019s model is as follows \u03c8(t) =W\u03c8\u03b4 ( d dt ) \u03b4(t) +Wyaw ( d dt ) d\u03c8(t), \u03d5(t) =W\u03d5\u03b4 ( d dt ) \u03b4(t) +Wroll ( d dt ) d\u03d5(t), (2) W\u03c8\u03b4(s) Wyaw(s)(Kdr +KvrWsway(s)), W\u03d5\u03b4(s) Wroll(s)(Kdp +KvpWsway(s)), where \u03c8(t), \u03d5(t), \u03b4(t) stand, respectively, for the heading (yaw), roll and rudder angles (see Fig. 1). The signals d\u03c8(t) and d\u03d5(t) are disturbances, describing the influence of the environment on the yaw moment and the roll moment respectively, Kdr,Kvr,Kdp,Kvp are the fixed parameters. The full stricture of the model is shown in Fig. 2. The disturbances acting on a marine craft are due to the wind, the waves and the current. The fast oscillations in the vessel\u2019s roll angle \u03d5(t) and its heading \u03c8(t) are mainly caused by the waves, whereas the current and the wind are changing much slower and their effect is usually modeled as a constant roll angle and stationary heading deviation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003439_s10854-020-04194-w-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003439_s10854-020-04194-w-Figure4-1.png", + "caption": "Fig. 4 Transformer model and vibration test points", + "texts": [ + " The graph shows the magnetization curve, where with DC bias it indicated by a solid line, while without DC bias it is shown by a dotted line. (a), (b), and (c) are represented, respectively, as the curve of magnetic flux, the magnetization curve of silicon steel, and the curve of the exciting current. By all appearances, DC bias causes distortion of the field current. It is a serious nonlinearity in the transformer magnetization curve. The 3D numerical model is implemented on a dry 5kVA transformer. It is the same as the one in vibration measurements. The physical model and its concerned points are as shown in Fig.\u00a04. This paper mainly discusses the effect of electromagnetic field on ferromagnetic materials. The sequential coupling method is selected in order to reduce computational complexity and time. As seen in Fig.\u00a05, COMSOL secondary development interface and hybrid programming are used to realize the numerical solution of electromagnetic field and force field of the transformer. The outer layer of the computer is transient time step iteration, while the inner layer is convergent iteration calculated sequentially by the physical field in each time point", + " It is because the deformation amplified from strain change of core through transformer structures. According to the simulation results, it is obvious that the core corners have the most serious effect with magnetic field circuit. So, the corner deformation could be an important concerning object for researching transformer vibration caused by silicon steel magnetostriction. In order to reflect the transformer vibration more accurately, the calculation waveforms of core and shell test points A, B, C, and D (marked in Fig.\u00a04) with and without DC bias are shown in Fig.\u00a08. The displacements in the vertical direction of point A and B on transformer are shown in Fig.\u00a08a, b. Figure\u00a08a shows the simulated result of point A under DC bias or not. In contrast to the normal work, the displacements increase and produce obvious distortion with 2A DC current. This would lead to odd components increasing in 1 3 vibration. Figure\u00a08b shows the simulated result of point B. The value of DC bias is distinctly different between with and without DC bias", + " The vibration acceleration of the transformer core is measured by piezoelectric acceleration sensors, the sensor\u2019s parameters are shown in Table\u00a01. The measured signal after amplification and conditioning is saved to the acquisition card. Then, industrial personal computers (IPC) analyze and process signals. The transformer which is the same with the simulation one is tested and its parameters are shown in Table\u00a02. And the experiment design and system are shown in Figs.\u00a09 and 10. During the test, the sensors are fixed on the core and shell. The locations of the test points are shown in Fig.\u00a04. Firstly, the no-load test method is used to measure vibration 1 3 1 3 under normal conditions. Secondly, the different DC current is input to the neutral point of the transformer. Finally, the test is implemented. Figure\u00a011 shows the vibration signals of core and shell at points A, B, C, and D under normal work and 2A DC current, respectively. By all appearances, the high-frequency components are obviously more than the simulation waveform in the experimental waveform. This can be explained by the noise of tests and structures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002756_1.5062894-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002756_1.5062894-Figure3-1.png", + "caption": "Figure 3: Dimensions of fatigue specimen, (a) undersized core for cladding, and (b) final dimensions.", + "texts": [ + " The gas atomized powder is spherical in shape with the size in the range of 15 - 45 m (Fig.2). Wrought L-605 flat substrate (127 mm x 50 mm x 10 mm) was purchased by NRC for laser cladding process development, while wrought L-605 round bars were supplied by UTC Aerospace Systems for making cylindrical fatigue specimens. All substrate coupons were heat treated in following conditions: Solution annealed at 1232 o C (2250 o F) for 1 hour with air quenching to ambient temperature. The wrought L-605 round bars were machined to undersized cylindrical cores for laser cladding (Fig.3a) and baseline fatigue specimens (Fig.3b), respectively. A Lasag Nd:YAG laser coupled to a fiber-optic processing head was used for laser cladding of L-605. The laser was operated in a pulse mode with an average power of 40 - 100 W. A Sulzer Metco 9MP powder feeder was used to deliver L-605 powder into the melt pool through a nozzle with a powder feed rate of 5 - 10 g/min. For laser cladding of fatigue coupons, the laser beam and the powder delivery nozzle were kept stationary, while the sample was rotated using a rotary table on a 5-axis CNC motion system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002619_iecon.2016.7793131-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002619_iecon.2016.7793131-Figure7-1.png", + "caption": "Fig. 7. Post-fault output voltage vector.(a) Vs located in sector I (b) Vs located in sector II.", + "texts": [ + " Other single switch under open-circuit fault can be analyzed in a similar way. In this section, a fault-tolerant control approach is proposed to improve the sine degree of stator current and to solve the problem of imbalance average output power in the two inverters under single switch open-circuit fault. To ensure the symmetrical structure of two inverters, when T11 is open, the switches T14 and T24 all keep the same state as shown in Fig. 6. According to the fault-tolerant topology in Fig. 6, the post-fault output voltage vector can be obtained as shown in Fig.7. When Vs is located in sector I and VI, inverter 2 synthesizes Vs alone. Inverter 1 is clamped to 000, as shown in Fig. 7(a). When Vs is located in sector III and IV, two inverters perform the opposite operation. It should be noted that when Vs is located in sector II, Vs can be modulated by OC from inverter 1 and OB from inverter 2, as shown in Fig. 7(b). When Vs is located in sector V, Vs can be modulated by OE from inverter 1 and OF from inverter 2. Table III shows the synthesis principle of reference voltage vector under single switch open-circuit fault in each sector. It can be seen from that this method can reduce the switching loss because the switching frequency of each inverter is decreased. The average output power of two inverters under fault tolerant control strategy proposed in this section can be given as follows: ) 3 4sin( 3 2 ) 3 2sin( 3 2 ) 3 5sin( 3 2 ) 3 sin( 3 2 )]3/5cos( )3/cos(cos2[ 6 )]3/4cos( )3/2cos(cos2[ 6 2 1 rs rs rs rs r rss s s r rss s s Vq Vp Vn Vm q pVI T TP n mVI T TP (7) where m, n, p and q denote the value of output voltage vector of inverter 1 and 2 respectively, Formula (7) can be simplified as: cos 2 1 cos 2 1 2 1 ss ss VIP VIP (8) It can be seen from (8) that P1=P2, so the average output power of two inverters can achieve balance under this faulttolerant control strategy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002293_jae-162048-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002293_jae-162048-Figure2-1.png", + "caption": "Fig. 2. Radial cross-sectional view.", + "texts": [ + " All rights reserved Stator yoke Mover yoke y z x Non-magnetic shaft 3-phase coils Permanent magnet Fig. 1. Structure of a helical teethed linear actuator. (Only stator is cross-sectional view). In this paper, a helical teethed linear actuator (HTLA) is proposed. This actuator has a unique structure with helical teeth and interior permanent magnets. The thrust force and torque of the HTLA are investigated by using 3-D finite element method (FEM). The performances under position and force control are simulated. The structure of the HTLA is shown in Fig. 1. Figure 2 shows the radial cross-sectional view (xy plane). The stator has a similar structure of rotary motors composed of a yoke and 3-phase coils. The shape of the tooth tip is helical. The mover consists of four sets of yokes with helical teeth, four sets of interior rectangular permanent magnets (NdFeB, Br = 1.3 T), and a non-magnetic shaft. The pitch of the helical teeth on the stator and the mover are the same. The magnetization direction of each magnet is shown in Fig. 2. The mover has 2 degrees of freedom for linear and rotary motions. Only the linear motion is used for the artificial muscle applications. The HTLA has a great advantage from the viewpoint of its productivity and accuracy against conventional cylindrical linear actuators because it does not require axially stacking permanent magnets. The structure of the HTLA without axially stacking permanent magnets has advantages against conventional cylindrical linear actuators. In the proposed structure composed of the helical teethed yokes with interior permanent magnets, the number of teeth can be increased without increasing the number of permanent magnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002174_j.engfailanal.2016.06.014-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002174_j.engfailanal.2016.06.014-Figure15-1.png", + "caption": "Figure 15: Residual hoop stress contour plots for S-type coach wheels at the end of one braking cycle consisting of 8 braking events for maximum heat input factors (a) one (synchronized braking) and (b) four.", + "texts": [ + " Second, we subject a wheel set to a braking cycle with maximum heat input factor 2 and then to a braking cycle with maximum heat input factor 4 (see Fig. 14). It can be seen from the figure that final wheel gauge change at end of second braking cycle and maximum gauge reduction, are essentially identical for both braking scenarios suggesting that these are primarily determined by maximum heat input factor. It can be seen that gauge reduction is lower in subsequent braking cycles when heat input factor for subsequent braking cycle is lower than the earlier one. Fig. 15 shows residual hoop stresses for S-type coach wheels at the end of a braking cycle for maximum heat input factors one (synchronized braking) and four. It can be seen from the figure that residual hoop stresses are compressive in tread region for synchronized braking but highly tensile for maximum heat input factor four (maximum hoop stress in this case is about 550 MPa). While thickness of tensile stresses for heat input factor four is around 25 mm, thickness of compressive hoop stresses for synchronized braking is around 50 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000533_tmag.2015.2442245-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000533_tmag.2015.2442245-Figure7-1.png", + "caption": "Fig. 7. Direction of rotational locomotion. (a) Clockwise. (b) Counterclockwise.", + "texts": [ + " Conversely, when a vibration component was displaced by amplitude A in the \u2212z-direction, a supporting force of 0.5 kA sin\u03b8 cos\u03b8 per vibration component acted on the pipe, as shown in Fig. 3. In the case of two vibration components, a supporting force of approximately 0.25 kA acted on the pipe. The frictional force between compound material A and the pipe wall oscillated, and the actuator moved while alternately sliding and stopping. As before, the displacements of vibration components 1 and 2 were synchronized during measurement. The actuator was held in the pipe, as shown in Fig. 6. Fig. 7(a) shows the displacements of vibration components 1, 2, and 5 when an electric current of 0.05 A was applied to electromagnets 1, 2, and 5, respectively, when the actuator was inserted into a pipe having an inner diameter of 35 mm. Fig. 7 shows the results when the displacements of vibration components 1, 2, and 5 were synchronized. Fig. 7(b) shows the results for the case in which there was a phase difference of 180\u00b0 between the vibration components. When the displacements of vibration components 1 and 2 were synchronized with the displacement of vibration component 5, the actuator rotated in a clockwise direction, as shown in Fig. 6. When the phase difference between these vibration components was 180\u00b0, the actuator rotated counterclockwise, as shown in Fig. 6. Fig. 8 shows the relationship between the phase difference of displacements for vibration components 1, 2, 5, and the rotational speeds when the inner diameter of the pipe was 35 and 40 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002539_iecon.2013.6699805-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002539_iecon.2013.6699805-Figure1-1.png", + "caption": "Fig. 1. The UPAT Tri-TiltRotor UAV platform in Hovering Flight", + "texts": [ + " Correspondingly, the UAV is desired to be designed around a multitude of flight characteristics, so as to combine the required flight capabilities, which are a) the navigation into constrained areas to perform shortrange tasks and b) the high-speed and prolonged endurance flight. The aforementioned goal characterisitc of operational duality can be satisfied with a convertible UAV design with a dual flight envelope and full actuation over the longitudinal\u2013 axis. Guided by the proposed principles, such a UAV design, shown in Figure 1, is being developed in the University of Patras (UPAT). This platform, namely the UPAT Tri\u2013TiltRotor (UPAT\u2013TTR) is equipped with rotor tilting mechanisms, and therefore is capable of achieving flight mode conversion from Vertical Take\u2013Off and Landing (VTOL) flying mode to Fixed\u2013 Wing (FW) mode. Moreover, the rotors\u2019 tilting mechanism and the corresponding authority of direct control of the projected thrust over the longitudinal axis leads to a vehicle that has full actuation on that axis even when operating around VTOL\u2013 mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000814_ssd.2014.6808791-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000814_ssd.2014.6808791-Figure2-1.png", + "caption": "Fig. 2: Tilt angles of the rotor w.r.t fixed frames", + "texts": [ + " When the rotors are aligned along the body z-axis, rotor 1 and rotor 2 are assumed to rotate counter-clock wise CCW, while rotor 3 and rotor 4 rotate clock-wise CWo The forward direction is taken arbitrary to be along the body x-axis. Assume that the rotational speed of the rotor i is given by Wi. Then we can say that the lifting thrust is given bw; and the drag moment is given by dw;. The orientation of the rotor is controlled by two rotations about the rotor fixed frame; ai, a rotation about the rotor y-axis, and f3i , about the rotor z-axis, as shown in Fig.2. To find the forces and torques generated by each tilted rotor on the air vehicle , let R\ufffd: be the rotational matrix of the rotor with respect to fixed axis at Oi. Since the axes at Oi are parallel to the body axes at the center of gravity of the air vehicle, then [O\ufffd 0 G (Jisai] R7,' = R\ufffd = 0 S{JiSai o Gai (1) 2 The thrust components of the ith rotor at the body CG. are then given by [0 0 G{JiSai] [ 0 ] Fi = 0 0 S{JiSai O2 o 0 Gai bWi (2) Similarly, the moments of a titled rotor consist of two parts, the drag moment, and the moments generated by the thrust components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001340_amr.1004-1005.1344-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001340_amr.1004-1005.1344-Figure2-1.png", + "caption": "Fig. 2 Contact area of axial ring rolling", + "texts": [ + " The profile curve of the contact surface is the intersecting line where roll surface and workpiece end face. In radial-axial ring rolling process, the contact surface is a part of the surface of the conical roller. S is the feed of the conical roller. In order to facilitate the calculation of the contact area, the contact surface can be approximated as plane because S is much smaller than R that is the outer diameter of the workpiece. Establish the rectangular coordinate system that the origin is the center of ring, as shown in Fig. 2, where O' is the conical point. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 141.117.125.76, Ryerson University Lib, Toronto-26/04/15,19:56:32) The plane equation can be expressed as cos ( )sinz x b S\u03b3 \u03b3\u2212 \u2212 = (1) Where, \u03b3 is the cone-top angle, r is the inside diameter of the ring, b is the distance from the center of the ring to the conical point. Ignore the elastic deformation of roll, equation of the surface of the conical roller can be expressed as 2 2 2 2( ) tan 0x b y z\u03b3\u2212 + \u2212 \u22c5 = (2) 2cos S Q R\u03b3 = \u22c5 (3) Q is the relative feed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure7-1.png", + "caption": "Fig. 7. The disc with crank arrangement.", + "texts": [ + " The connecting rod is the link that connects the bump to the crank. It has holes on both ends for mounting it on the bump and on the disc crank using M12 bolts. The dimensions proposed have been used to calculate the critical buckling load Fcr in Sec- Please cite this article as: A. Sinha, S. Mittal, A. Jakhmola et al., Green energy ge ings, https://doi.org/10.1016/j.matpr.2020.06.248 tion 2.2. It will be made of 5 mm thick IS2062 Grade E250 sheets. It is shown in Fig. 6. The disc crank is a crank in the form of a disc as shown in Fig. 7. It has a hole at a distance of 5 cm from the center for mounting the connecting rod using an M12 bolt. It is itself mounted on a shaft of outer diameter 25.4 mm. The diameter of the disc is 15 cm. A disc shaped crank has been used because it is easy to design and manufacture. It will be made of 5 mm thick IS2062 Grade E250 sheets. A compound gear train has been designed and assembled to serve as the power transmission and speed multiplication unit. Spur gears have been used with specifications and dimensions as mentioned in Section 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure5-1.png", + "caption": "Figure 5. Bodies i and j connected by spherical joint.", + "texts": [ + " To D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 remove the redundant constraints, one pin joint is made a spherical joint and another pin joint is made the parallel primitive joint as shown in Figure 4 [13,14]. The spherical joint removes three relative translations along the joint x, y and z axes, while the parallel primitive joint removes two relative rotations about the joint z and x axes. The combination of both spherical and parallel joints results in an equivalent revolute joint between two crawler shoes. The bodies, labelled 2\u201314, are connected by revolute joints at the locations of the first link pin. In Figure 5, the points Pi and Pj defined along the joint axis on bodies i and j are always coincident during the entire motion of the crawler track [15]. The three constraint equations for spherical joints from Shabana [15] can be derived from Equation (13). KS ui; uj \u00bc rPi rPj \u00bc 0 (13) i = 2, 3, \u2026, 13 and j = 3, 4, \u2026, 14 and rPi , rPj are global position of point P on body i and body j. Equation (13) can be rewritten in terms of generalised coordinates of bodies i and j defined by Equation (14). KS ui; uj \u00bc rPi rPj \u00bc Bi \u00fe Ais 0 Pi Bj Ajs 0 Pj \u00bc 0 (14) In Equation (14), Ai and Aj are the rotation matrices that define the orientation of the body i and j with respect to the global coordinate system [15], and s0Pi and s0Pj are the positions of the point P on body i and body j with respect to its centroidal coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003402_s11071-020-05852-8-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003402_s11071-020-05852-8-Figure5-1.png", + "caption": "Fig. 5 Beam NES schematic idealization", + "texts": [ + " The system is idealized as massless bar of length l, pivoted at point A with a concentrated mass m at the tip (point B), and constrained by a spring force fsp of nonlinear characteristics, given by the general expression inEq. (7), and a damping force fd of viscous characteristics. The pivot point is fixed to a moving cart representing the host structure base acceleration u\u0308b. The motion of the bar is defined completely by its rotation angle from the vertical equilibrium position, \u03c6. This system is shown schematically in Fig. 5. Several assumptions need to be made, namely that: (1) the beam is rigid, so all bending characteristics from its elastic physical features are assumed to be captured by the nonlinear spring characteristic; (2) the radius of curvature of the path described by the mass is constant, which is not true in reality as the length of the beam changes progressively as it wraps around the curved surface of the boundary; and, (3) the beam is massless, as its actual mass is much smaller than the concentrated mass at its tip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003421_aim43001.2020.9158930-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003421_aim43001.2020.9158930-Figure7-1.png", + "caption": "Fig. 7. (a) Gripper model with a grasp frame, (b) Gripper model used to grasp object.", + "texts": [ + " The graspable outlines produce grasp candidates subspace G\u0303 wherein a grasp candidate g(i, x) \u2208 G\u0303, i \u2208 N indicates the index of the grasp candidate line, and x \u2208 R is the position of the grasp candidate on the line. The proposed algorithm was used to generate rigid grasp pose candidates for the IKEA Stefan chair. The source codes are available on Github at https://github.com/psh117/fgpg. Figure 8 (a) depicts the result of the proposed method. The first row illustrates the subspace of the grasp candidates. The second row visualizes the candidates using the 3D gripper model shown in Fig. 7. The third row shows enlargements of the red boxes of the second row. Each column illustrates the back part, seat part, apron part, and a side part of the IKEA Stefan chair. Figure 8 (b) depicts the candidate generation results using antipodal-based method [7] in the same order. The first row illustrates the preliminary points. Figure 9 shows the average insertion depth of the grippers shown in Fig. 8. It is evident that the grasp candidates depicted in Fig. 8 (b) can barely utilize the palm of the gripper whereas the grasp candidates represented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure5.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure5.2-1.png", + "caption": "Figure 5.2 Revolute, prismatic, and spherical joints.", + "texts": [ + " Here and hereafter, the subscript indices i, j, and k are ordered so that ijk \u2208 {123, 231, 312}. If ab is a revolute joint, the relative orientation of ba with respect ab is described by one joint variable, which is the angle \ud835\udf03ab that rotates u\u20d7(ab) i and u\u20d7(ab) j , respectively, into u\u20d7(ba) i and u\u20d7(ba) j about n\u20d7ab = u\u20d7(ab) k = u\u20d7(ba) k . The relative location of ba with respect to ab does not change. In other words, the distance dab = OabOba remains constant. Usually, a revolute joint can be arranged so that dab = 0. A revolute joint is illustrated in Figure 5.2 with its mating kinematic elements. If ab is a prismatic joint, the relative location of ba with respect to ab is described by one joint variable, which is the sliding distance sab = OabOba along n\u20d7ab = u\u20d7(ab) k = u\u20d7(ba) k . The relative orientation of ba with respect to ab does not change. In other words, the angle \ud835\udeffab remains constant. This angle describes the constant orientation of u\u20d7(ba) i or u\u20d7(ba) j with respect to u\u20d7(ab) i or u\u20d7(ab) j . Usually, a prismatic joint can be arranged so that \ud835\udeffab = 0. A prismatic joint is illustrated in Figure 5.2 with its mating kinematic elements. Another joint, which is encountered quite frequently, is a spherical joint. Its mobility is \ud835\udf07ab = 3. At a spherical joint, the kinematic elements ab and ba share Oab = Oba as their common center. Therefore, the relative location of ba with respect to ab does not change. On the other hand, since \ud835\udf07ab = 3, the relative orientation of ba with respect to ab is described by three independent joint variables, which are usually selected as the Euler angles (\ud835\udf19ab, \ud835\udf03ab, \ud835\udf13ab) of an appropriate sequence. The appropriate sequence for a particular spherical joint is case dependent and it is selected so that it helps simplify the expressions of the relevant kinematic relationships. A spherical joint is illustrated in Figure 5.2 with its mating kinematic elements. The appropriate sequences selected for two different spherical joints can be seen as demonstrated for a spatial slider-crank mechanism studied as an example in Section 5.6. A large collection of joints will be considered later in Chapter 6 with all the details of their kinematic descriptions and the relevant equations. Consider a kinematic chain that consists of the rigid bodies o, a, b, ..., z. In this kinematic chain, o is indicated as the leading member because it is selected as the base (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001089_s00170-015-7021-6-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001089_s00170-015-7021-6-Figure12-1.png", + "caption": "Fig. 12 Cutting region when peripheral cutting edge takes part in machining and yB<>: c \u00bc arccos yB r d \u00bc arccos yC r 8<: c 0 \u00bc arccos yC r d 0 \u00bc arccos yB r 8<: a 0 \u00bc arccos yB r b 0 \u00bc arccos yA r 8<: \u00f038\u00de ( 5 ) F o r I E i i n p a r t V, w h e r e ffiffiffiffiffiffiffiffiffiffiffiffiffiffi h\u2212Rc\u00f0 \u00dep 2 \u00fe yB 2 < r < R2; si \u00bc s2 \u00fe s2 0 \u00fe s1 c \u00bc arcsin h\u2212Rc r d \u00bc arccos yC r 8><>: c 0 \u00bc arccos yC r d 0 \u00bc arccos yB r 8<: a \u00bc arccos yB r b \u00bc arccos yA r 8<: \u00f039\u00de (6) For IEi in part VI, where R2R1, the cutting region can be divided into four different parts, as shown in Fig. 13. In this case, part I, part II, and part III are all the same as they are in Fig. 11. So the corresponding swept area si of the three parts can be calculated by Eqs. (31), (37), and (33). While part IV can be treated the same as part IV in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure20-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure20-1.png", + "caption": "FIG. 20. Deformed cross-sections for the (a) straight crease and (b) bending crease deformation modes.", + "texts": [ + " Instead, the indenter contacts the creases some distance from vertex. Additionally, the analytical models assume that the disk is supported where the creases reach the edge of the disk; however, for the tests the supports are located closer to the center of the disk. Using the geometry of the disk and test apparatus, a conversion is made between the displacement of the vertex obtained from the models and the corresponding experimentally measured displacement of the indenter. For the straight crease, shown in Fig. 20(a), \u03b4 is the model displacement measuring the distance from the central vertex to where the crease reaches the disk edge. The experimental displacement, \u03b4\u2217 1 , is measured from the supports to the contact point of the indenter and the creases. Using the geometry of the test, the relationship between the experimental and model displacements is \u03b4\u2217 1 = \u03b4 \u2212 di 2 ( \u03b4 R ) \u2212 \u03b4 R \u23a1 \u23a3R \u2212 L 2 \u221a 1 \u2212 ( \u03b4 R )2 \u23a4 \u23a6 \u2248 \u03b4(L \u2212 di ) 2R . (B1) For the bending crease a dimple of radius , measured along the crease, forms and deformation primarily occurs within this dimple; see Fig. 20(b). Similar to the straight crease, \u03b4 is the model displacement coordinate measuring the distance from the central vertex to the dimple edge. The experimental displacement, \u03b4\u2217 1 , is measured from the supports to the contact point of the indenter and the creases. Using the geometry of this configuration, the relationship between experimental and model displacements found: \u03b4\u2217 2 = \u03b4 \u2212 di 2 ( \u03b4 ) = \u03b4 ( 1 \u2212 di 2 ) . (B2) When the number of creases is greater than two, the creases do not initially lie in the same plane as the supports", + " When the deformation of the disk raises the central vertex above the supports (\u03b4 < 0), the contact between the indenter and the disk is at the hole edge, therefore \u03b4\u2217 3 = ( L 2 tan \u2212 r sin ) \u2212 ( L 2 tan 0 \u2212 r sin 0 ) , where the first bracketed term is the distance from the support to the edge of the hole in the deformed state and the second bracketed term is the distance from the support to the hole edge in the natural state. Noting that sin = \u03b4/R and 0 is the crease inclination in the natural state, \u239b \u239e N 2 3 4 5 0 0 \u22120.12 \u03b2 \u22120.17 \u03b2 \u22120.19 \u03b2 043001-16 When the creases pass through the plane of the supports (\u03b4 > 0) the contact point switches to the indenter edge, as shown in Fig. 20: ( \u03b4\u2217 3 ) \u03b4>0 \u2248 (L \u2212 di ) \u03b4 2R \u2212 L 2 tan 0 + r sin 0. (B4) Noting the first term in Eq. (B4) is equal to \u03b4\u2217 1 , by analogy for the bending crease model, \u03b4\u2217 4 = r sin 0 \u2212 L 2 tan 0 + {( 1 \u2212 r ) \u03b4 if \u03b4 < 0,( 1 \u2212 di 2 ) \u03b4 if \u03b4 > 0. From Lechenault and Adda-Bedia [14] the crease inclination in the natural states are given in Table II. [1] L. Wilson, S. Pellegrino, and R. Danner, in Proceedings of the 54th AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference (AIAA, Boston, MA, 2013)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002324_jae-162080-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002324_jae-162080-Figure1-1.png", + "caption": "Fig. 1. Rolling element bearing\u2019s structure.", + "texts": [ + " The outline of this paper is listed as follows: Section 2 presents signal characteristics of rolling bearings faults; Section 3 discusses the theory of WPT and the performance comparison of three networks; damage diagnosis procedure is proposed in Section 4; the simulation of bearing damage diagnosis and conclusion are presented in Section 5 and Section 6. Since that rolling bearings are usually used, some parts of rolling bearings are easy damaged [9]. Their main structure and load distribution can be seen clearly in Fig. 1. In spite of various damage forms, bearing failures always appear on the surface of inner race, outer race, and balls in the early stage. Faults in different locations will produce unique frequency components in the machine vibration signals. Generally, the outer race is stationary. Some fault frequency can be expressed as follows: FO = 1/2 \u00b7NbFR(1\u2212Db cos \u03b8/Dc), (1) FI = 1/2 \u00b7NbFR(1 +Db cos \u03b8/Dc), (2) FB = Dc/Db \u00b7 FR[1\u2212 (Db cos \u03b8/Dc) 2], (3) where FR is the rotor mechanical frequency, and FI , FO, FB are fault frequency of inner raceway, outer raceway and ball" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001502_s0263574714002276-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001502_s0263574714002276-Figure1-1.png", + "caption": "Fig. 1. Prototype of a 6-DOF Stewart platform. (a) Side view; (b) vertical view.", + "texts": [ + " A simple velocity estimation strategy is introduced in Section III. The main results are presented in Section IV, where a novel adaptive vector SMC scheme based on position measurements is proposed for the trajectory tracking of the Stewart platform. Section V discusses the implementation aspects, computational efficiency, and the chattering phenomenon of the proposed sliding mode controller. The simulation results are presented in Section VI, and the conclusions in Section VII. The prototype of a Stewart platform, as shown in Fig. 1, consists of six extendable actuators connecting a movable platform to a fixed base with spherical and universal joints. Ob \u2212 XbYbZb, OP \u2212 XPYPZP, da, db, ra, rb, and l0 represent the reference frame, platform frame, upper-joint spacing, lower-joint spacing, upper-joint radius, lower-joint radius, and middle-actuator length, respectively. Moreover, the generalized coordinate vector describing the position and orientation of the movable platform is denoted \u03c7= [x, y, z, \u03c6, \u03b8, \u03c8]T (which is the Euler angle description)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003065_s40684-020-00217-3-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003065_s40684-020-00217-3-Figure5-1.png", + "caption": "Fig. 5 Thermal analysis results of the turbine blade by laser preheating", + "texts": [ + " The governing equation of the 3D transient analysis is represented by Eq. (1), as follows: where is the thermal diffusivity and T , Q , t , , C and k are the temperature, heat generation rate, time, density, specific heat and thermal conductivity, respectively. The initial condition at time t = 0 is given by Eq. (2) as The boundary conditions can be described by Eq. (3) as where T , q , h , and T0 are the surface temperature, heat flux, heat transfer coefficient and ambient temperature, respectively [23, 26]. Figure\u00a05 presents the results of the thermal analysis conducted to determine the cutting depth for LAMill based on the proper preheating temperature (600 \u2103). It was found that the maximum surface temperature of the turbine blade was 598.61 \u2103. A sectional view, which assisted in determining the cutting depth for the LAMill process, found a value of approximately 1 mm. The implies a depth of up to 400 \u2103, where the tensile strength is reduced by more than half in the direction of the cutting depth as compared to the tensile strength at room temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002293_jae-162048-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002293_jae-162048-Figure5-1.png", + "caption": "Fig. 5. System configuration of the potision control.", + "texts": [ + " Moreover, the kinetic machine model includes coulomb and viscous friction models. The coulomb frictions in the linear and rotary directions were 1 N and 1 mNm, respectively. The viscous frictions in the linear and rotary directions were 1 Ns/m and 1 mNms/rad, respectively. The performances of the actuator under position control were computed when the mover moved from 0 to 10 mm. The maximum velocity and acceleration of the target trajectory are 0.5 m/s and 1 m/s2, respectively. The system configuration of the position control is shown in Fig. 5, which contains a PID controller. An input value of the PID controller is a distance between the target and current positions. An output value of the PID controller is the q-axis current. The d-axis current is always set to 0. The control frequency of the PID controller is 1 kHz. The result is shown in Fig. 6. The P gain was 30, the I gain was 0.05, and the D gain was 0.001. Mass of the mover was 212 g. The calculated and target positions show a good agreement and this actuator shows high response characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001184_ijaac.2013.057058-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001184_ijaac.2013.057058-Figure1-1.png", + "caption": "Figure 1 Structure of DFNN", + "texts": [ + " To solve these problems, we propose an adaptive dynamic fuzzy neural network-based decouple sliding mode control algorithm using DFNN and fractional PI control law in Section 5. First, a brief introduction of dynamic neural network controller is given below. In this paper, DFNN is constructed to approximate the non-linear functions in the equivalent control term of DSMC. It invokes DFNNs to represent the unknown dynamics of uncertain non-linear system. A detailed study of the DFNN learning algorithm can be found in the work of Wu and Er (2000) and Er et al. (2005). The DFNN architecture is shown in Figure 1. It consists of five layers. Detailed descriptions of DFNN can be summarised as: \u2022 Layer 1: Input-variable layer. Each node in this layer represents an input linguistic variable. This is the layer where the input signals first enter the neural network. \u2022 Layer 2: Each node in this layer represents a membership function (MF) which is in the form of Gaussian functions: ( ) ( )2 2 , ( 1,2, , ; 1, 2, , ), i ij j x c ij ix e i n j r\u03c3\u03bc \u2212 \u2212 = = =\u2026 \u2026 (18) where n is the number of input variables, r is the number of membership functions, \u03bcij(xi) is the membership function of xi, cij is the centre of the jth Gaussian membership function of xi, and \u03c3j is the width of the jth Gaussian membership function", + " \u2022 Layer 4: This layer consists of r (normalised) nodes. The output of the jth node is given by 1 \u02c6 . j j r k k R R \u03bc = = \u2211 (20) \u2022 Layer 5: Each node in this layer represents an output variable as the summation of incoming signals. Each of which is weighted according to equation (21). This layer performs defuzzification (weighted average) of the output as follows: 1 \u02c6\u02c6( ) ( ) ( ), r j j j f x x f x\u03bc = =\u2211 (21) where \u02c6 j\u03bc is the normalised output of layer 4\u2019s jth node, f(x) is the value of an output variable and \u02c6 jf is the weight of jth rule. In Figure 1, NNj represents a simple NN, for the TSK model, the output of NNj can be represented as 0 1 1 \u02c6 ,j j j jn nf x x\u03b8 \u03b8 \u03b8= + + + (22) where \u03b8ji (i = 0, 1 \u2026, n) are the adjustable real-valued parameters. Using a simplified notation, by applying TSK fuzzy inference mechanism and considering 2\u0302 ( )f x as an output, (21) can be rewritten in a compact form: ( )2 2 22\u0302 ( ),T f f ff x x\u03b8 \u03b8 \u03be= (23) by similar procedure, the function g2(x) can be modelled as follows: ( )2 2 22\u02c6 ( ),T g g gg x x\u03b8 \u03b8 \u03be= (24) In the network construction step, first we need to give an error criteria of neuron generation, let the criterion be defined as , 1, 2, , ,j j je t y j r= \u2212 = (25) where r is the number of existing RBF units" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000953_oceans.2014.7002998-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000953_oceans.2014.7002998-Figure1-1.png", + "caption": "Fig. 1. System configuration of the 7-function manipulator", + "texts": [ + " THE 7-FUNCTION MANIPULATOR SYSTEM DESIGN The 7-function manipulator system could be mounted on remotely-operated vehicles or human occupied vehicles to perform underwater tasks, and normally works in a masterslave mode. The system is divided into surface part and underwater part, the former is composed of a surface control 978-1-4799-4918-2/14/$31.00 \u00a92014 IEEE box and a power supply junction box, while the latter is composed of a hydraulic control valve pack, a slave arm, a hydraulic power unit and a pressure compensator, and linked to the surface part with a cable. The whole system configuration is shown in Fig. 1 and the system picture is shown as Fig. 2. The slave arm is the system\u2019s actual actuator which is powered by hydraulic with 6-DOFs movement joints and a gripper, each joint equipped with a displacement sensor to achieve servo control. The slave arm is controlled by the master arm, besides automatic extension and recovery functions are included[14,15]. Surface control box is the control terminal of the system, consisting of a master controller, a control panel and a master arm. Operators can choose control mode and set parameters of the system through the function keys on the control panel, an LED monitor is used for visualized human-machine interaction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003167_physreve.101.052410-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003167_physreve.101.052410-Figure4-1.png", + "caption": "FIG. 4. The schematic view of cell behavior and push due to heterophilic adhesion for (a) large ha and for (b) small ha. (c) Magnified view of the leading edge. Solid and dashed arrows in (a) and (b) express the polarity and net motion, The solid arrow in (c) represents the push direction of the transported cell on the scaffold.", + "texts": [ + " These states are characterized by the net direction of the polarity P. From Eq. (5), these directions of P follow the direction of dRm/dt averaged over cells. Therefore, the directions of P indicate that there exists a net collective cell motion in the x direction for small ha and in the y direction for large ha. Therefore, to explain these states, we consider the direction of the cell motion in these states. For large ha, the rapid motion of transported cells generally tends to occur in the y direction, as shown in Fig. 4(a). This is because the transported cells are restricted by the scaffolding tissue and thereby mainly aligned their motion in the y direction. In this case, because rapid motion due to large ha exhibits an order state as a polarity memory effect [40,43\u201347], the net motion of the transported cells aligns P in the y direction through Eq. (5). In contrast, for small ha, the x-directional motion of transported cells implies that the restriction of scaffolding tissue is apparently ineffective in the determination of P. This is due to ha being too small to induce order in cell motion in the y direction as shown in Fig. 4(b). The contribution of y-directional motion cancels itself and is averaged out at a low value in the determination of P. Instead, the small net motion of the transported cells in the x direction, which is negligible for large ha in comparison with that in the y direction, emerges. This net motion is driven by pushing due to the energy gain of the heterophilic adhesion between the transported cells and the scaffolding tissue as shown in Fig. 4(c). This net motion finally aligns P in the x direction through Eq. (5). As a result, the two states appear to be dependent on the values of ha. Finally, we discuss the possibility of the realization of the state for small ha. In real systems, scaffolding tissues are supposed to be fixed in space. In this case, because of the absence of net motion in the x direction, the motion of the transported cells is expected to exhibit a simple random walk in the y direction. As a result, the transport of cells is expected to become diffusive for small cell adhesion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001539_ssp.220-221.37-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001539_ssp.220-221.37-Figure1-1.png", + "caption": "Fig. 1. Example of mesh generated for bearing with \u03bb = 0.5", + "texts": [ + " This research contains results of hydrodynamic pressure distributions for slide journal bearing of dimensionless length L1 = 0.5 and relative bearing clearance \u03c8 = 0.004. The rotational speed of bearing journal \u03c9 = 300 rad/s (2865 rpm). The simulations concern bearing lubricated with oil which has viscosity properties as Shell Helix Ultra AV-L, investigated in the paper [3]. The investigated bearing geometry was created using ANSYS DesignModeler software. Mesh was generated with ANSYS Workbench Meshing software. Figure 1 shows mesh with 9210 elements and 12992 nodes generated for bearing with relative eccentricity [5] \u03bb = 0.5. Mesh was generated for full wrap angle. The oil film wedge was sliced into three layers with respect to the radial component. Solution was obtained with ANSYS Fluent. The solution methods, equations etc., are widely described in [6]. Figure 2 shows the hydrodynamic pressure distribution for such bearing with \u03bb = 0.5, assuming that the viscosity of lubricating oil depends on temperature, as Shell Helix Ultra AV-L at temperature 90 degrees of Celsius" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002056_1350650116631453-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002056_1350650116631453-Figure1-1.png", + "caption": "Figure 1. A schematic diagram of a cylinder roller bearing.", + "texts": [ + " Equations are solved by Euler method with fixed time step. A mixed elastohydrodynamic model considering surface roughness and finite-length contact is established to study the lubrication performance in interface between the maximum loaded roller and inner race. Taking the angular velocities obtained from a dynamic model as an input, the effect of acceleration on lubrication characteristics is analyzed. Mathematic formulation of dynamic model The schematic of cylinder roller bearing is given in Figure 1. The inner race rotates with shaft (angular velocity !) while the outer race is fixed. !j represents angular velocity of rollers about its axis and !c represents angular velocity of cage. Ri is the radius of inner race, Ro is the radius of outer race, Rm is the pitch radius of bearing, Rr is the radius of rollers. A radial force W is loaded on bearing. According to load balance (Figure 1), force balance equation is written as25,26 W \u00bc Q0 \u00fe 2 X Q cos \u00f01\u00de at UNIV CALIFORNIA SAN DIEGO on February 24, 2016pij.sagepub.comDownloaded from whereW is bearing load, Q0 are the maximum load on roller ( \u00bc 0), is the position angle, and Q is the load distribution at the different angle position. According to conformability of deformation, the deformation can be described as25 \u00bc max cos \u00f02\u00de In accordance with the relationship between the deformation and the contact load, one equation can be attained as25 Q Q0 \u00bc max 1=t \u00f03\u00de where t is the coefficient, t\u00bc 3/2 for ball bearings, t\u00bc 10/9 for roller bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000054_iccia49288.2019.9030886-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000054_iccia49288.2019.9030886-Figure3-1.png", + "caption": "Figure 3. The schematic of the power transmission systems.", + "texts": [ + " To derive the dynamics equations, let us neglect the flexibility of cables, and write the dynamic model of DSCR in the following general form [22]: M(x)x\u0308+C(x, x\u0307)x\u0307+G = F =\u2212JT \u03c4, (4) where M, C and G respectively denote the mass , Coriolis and centrifugal matrices and the gravity vector. Furthermore, x and F are vectors denoting the generalized coordinates of the end-effector position and the Cartesian wrench applied to it. Moreover, J denotes the Jacobian matrix and \u03c4 the cable forces vector. On the other hand, the dynamics of the actuators and power transmission systems shown in figure 3, are represented by IM \u03b8\u0308 M +DM \u03b8\u0307 M \u2212 rM\u03c4P1 = rMu, (5) Furthermore, the dynamic equations for the power transmission system can be written as: IPi \u03b8\u0308 Pi +DPi \u03b8\u0307 Pi \u2212 rPi\u03c4Pi+1 =\u2212rPi \u03c4Pi i = 1 : n\u22121. (6) For the last pulley we have: IPn \u03b8\u0308 Pn +DPn \u03b8\u0307 Pn \u2212 rPn \u03c4 =\u2212rPn\u03c4Pn , (7) where, n denotes the number of pulleys, \u03b8 M , \u03b8 Pi the angle vectors of drum and ith pulley, Im, IPi the inertia matrix of drum and ith pulley, DM , DPi the viscous friction matrix of drum and ith pulley and rM , rPi , the radius of drum and pulley" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000834_12.2083777-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000834_12.2083777-Figure11-1.png", + "caption": "Figure 11: The in-house developed 2D - polygon scanning device. Due to the constructional concept it can be easily implemented in any setup as a galvanometer scanner replacement. The exchangeable f-theta optic allows the adaption of the spot diameter and scan field size to the process. (b) Precisely rastered scan field generated above the building platform by the pilot laser at 1000 ms-1 for the fast axis. The deviation in the slow axis is driven by a conventional galvanometer.", + "texts": [ + " In this manner, by assuming that the coating speed is not the limiting factor and, further, due to the disproportional fast generating of the melt depth at thin powder layers [14], now this kind of the developed laser micro sintering has the potential to achieve more than ten times faster build-up rates than conventional laser sintering/melting technologies. However, this can only be achieved by irradiating high laser power, such as supplied by high-intense (high brilliant) high power cw lasers. For example, the above mentioned 100 ns interaction time of the laser beam with the powder layer can be achieved by deflecting a high brilliant cw laser beam of 20 \u00b5m spot diameter at 200 m/s scan speed. This high scan speed was achieved in our experimental study using an inhouse developed 2D polygon scan system, as shown in Fig. 11. Proc. of SPIE Vol. 9353 93530P-7 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 10/26/2015 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx The micro part presented in Fig. 12 is the first laser micro sintered 3D specimen produced using mechanical high powder compaction in combination with high-intense cw laser irradiation. In this approach, relatively thick sinter layers of 10 \u00b5m and a scan speed of 12 ms-1 were applied. Although the applicable scan speed was limited by the thick powder layer (it was reduced to wet the underlying sinter layer), considerably higher building rates were achieved than those reported previously for micro sintering with merely 1 \u00b5m thick sinter layers and a maximum scan velocity of 1 ms-1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001400_detc2013-12646-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001400_detc2013-12646-Figure1-1.png", + "caption": "FIGURE 1. THE CABLE-DRIVEN ROBOT AND A GRIPPER AS ITS END-EFFECTOR", + "texts": [ + " Force-Feasible Workspace (FFW) is presented in [4] for point-mass cable robots. This workspace is defined as all positions where the end-effector is able to exert the desired force set to its surroundings having positive bounded magnitudes of tensions in all cables. Also Wrench-Feasible Workspace (WFW) is studied in [5]. In this paper the workspace of a cable-driven robot, previously proposed by authors, is investigated. Our proposed cable robot in [6], uses a gripper as its end-effector which makes it suitable for object handling as shown in Fig. 1. As it is explained in the following sections, this cable robot is actuated by three active cables, which control the position of the endeffector, during positioning operation and it becomes a redundant robot as the passive cables are locked at end points of positioning path, where the grasping operation accomplishes. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/02/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2013 by ASME In [6], force feasible workspace of this robot is investigated during positioning operation and the formulation of obtaining FFW for the proposed robot is presented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001129_iraniancee.2013.6599611-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001129_iraniancee.2013.6599611-Figure5-1.png", + "caption": "Fig. 5: Adaptive control of the load angle of the systems with known and the estimated parameters", + "texts": [ + " Table II shows the results of this identification scheme. As it can be seen, very good estimate of system parameters are obtained. To show the capabilities of the Indirect Adaptive Control method, a simulation for the sample two-mass is performed when its parameters are unknown. The initial parameters for the Adaptive Control are the same as identification simulation as shown in tables I, II. The control object is to regulate the load angle to 5 radians and the system starts from 0 radian initial condition. As shown in Fig. 5, in spite of not knowing the system parameters, the application of control law (47) together with identification schemes results very good performance. The load angle for controlling the two-mass system when the parameters are known is also shown in Fig. 5. The transient difference between these two cases is due to the simultaneous execution of the identification schemes together with the indirect adaptive algorithm. 6. Conclusions Two-mass mechanical systems with backlash nonlinearity and with unknown physical parameters are studied in this paper. The main goal is to design an adaptive position controller for this kind of systems. To achieve this goal, linear in parameters models are derived from the system model and a two-stage identification procedure is proposed for estimating the unknown parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002593_978-3-319-46155-7_6-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002593_978-3-319-46155-7_6-Figure10-1.png", + "caption": "Fig. 10 Final lightweight part manufactured by SLM", + "texts": [ + " The objective criterion of the optimization is a volume fraction of 15% of the design space. The part is optimized regarding stiffness. Figure 9 shows the optimization result as a mesh structure and an FEM analyzes for verification of the structure. The maximum stress is approx. 300 MPa, which is below the limit of Yield Strength of 410 MPa. Before manufacturing, the surfaces of the optimized part are smoothened to improve the optical appearance of the part. Compared to a conventional part (90 g), a weight reduction of approx. 15% (final weight 77 g, see Fig. 10) was achieved. Further improvements to increase the productivity of the process are needed for series part production. The almost unlimited design freedom offered by SLM provides new opportunities in lightweight design through lattice structures. Due to unique properties of lattice structures (good stiffness to weight ratio, great energy absorption, etc.) and their low volume, the integration of functional adapted lattice structures in functional parts is a promising approach for using the full technology potential of SLM (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003073_iet-epa.2019.0941-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003073_iet-epa.2019.0941-Figure4-1.png", + "caption": "Fig. 4 Structure of WT-IPMSM (a) Operation trajectory, (b) Vector diagram", + "texts": [ + " The weakening of air-gap magnetic field IET Electr. Power Appl., 2020, Vol. 14 Iss. 7, pp. 1186-1195 \u00a9 The Institution of Engineering and Technology 2020 1187 Authorized licensed use limited to: University of Massachusetts Amherst. Downloaded on July 05,2020 at 09:26:27 UTC from IEEE Xplore. Restrictions apply. reduces stator core loss density, thereby the temperature gradient distribution and operation reliability of IPMSM have been adjusted and improved. The control strategy of 45 kW EV WT-IPMSM is shown in Fig. 4. From Fig. 4a, the maximum torque per ampere (MTPA) is applied in region I and the running track is O-H-F-D-A. In this region, the WT-IPMSM operates at constant torque along the curve, so the corresponding speed is maximum speed and point A is the position of maximum output torque. Similarly, the MTPA is still applied in region II with O-H-F-D track and in region III with O-H-F track. The region IV is controlled by the maximum torque per voltage (MTPV). When the running track arrives at point F in region III, it needs to continuously increase speed. Hence, the running track is F-G-I-C. Besides, the speed at point C is positive infinity, so it is considered to be the ideal operating state. Furthermore, the application of WT strategy to sub-region operation still follows the vector control method, so it needs to satisfy voltage limit ellipse and current limit circle [1]. The vector diagram corresponding to (1) and (2) is shown in Fig. 4b. Equations (3) and (4) are voltage and current limit constraint, respectively. Meanwhile, (5) shows the electromagnetic torque including PM torque and reluctance torque. ud = Rsid \u2212 \u03c9rLqiq (1) uq = Rsiq + \u03c9rLdid + \u03c9r\u03c8f (2) ud 2 + uq 2 \u2264 us max 2 (3) id2 + iq2 \u2264 is max 2 (4) Te = p[\u03c8fiscos \u03b3 + 1/2(Ld \u2212 Lq)is2sin(2\u03b3)] (5) where Ld and Lq are the inductances of the d\u2013q-axis, ud and uq are the stator voltages of the d\u2013q-axis, id and iq are the stator currents of the d\u2013q-axis, |us|max and |is|max are the maximum values of terminal voltage and phase current, is is the stator current vector, Rs is the phase resistance, \u03c9r is the rotor velocity, p is the number of pole pairs, \u03b3 is the angle difference between is and E0, Te is the electromagnetic torque, and \u03c8f is the PM flux linkage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000589_aeat-12-2012-0258-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000589_aeat-12-2012-0258-Figure1-1.png", + "caption": "Figure 1 yaw \u2013 pitch gimbal configuration", + "texts": [ + " Information regarding the fact that state feedback controllers are synthesized through convex optimization and robust controllers are synthesized through mixed sensitivity and model matching approaches has been described. Simulation results, in which tracking performances of the control channels are compared and analyzed are presented next and the ultimate impacts of disturbance torque inputs on LOS are examined. The conclusions of the paper follow this section. A sketch of a typical yaw \u2013 pitch gimbal assembly with three defined coordinates is given in Figure 1, where, coordinate B: xB yB zB is fixed to the platform, coordinate O: xO yO zO is fixed to the out frame and coordinate I: xI yI zI is fixed to the inner frame. Rotation of the outer and inner frames occurs about the xO and yI axes, respectively, and the respective gimbal angles are designated and . According to the torque balance relationships for the inner and outer frames, there exist equations (1) and (2), JI\u0307I I JI I TIx TP TIz TIm 0 TIfc 0 (1) JO\u0307O O JO O TY TOy TOz Om TOfc 0 0 TIx TP TIz 0 TIfc 0 (2) where, JI is the inertia matrix of the inner gimbal, with JI diag JIxx, JIyy, JIzz ; JO is the inertia matrix of the outer gimbal, with JO diag JOxx, JOyy, JOzz ; B is the platform angular rate in relation to inertial space about coordinate B; O is the outer gimbal angular rate in relation to inertial space about coordinate O; I is the inner gimbal angular rate in relation to inertial space about coordinate I; TIm is the inertial torque induced by unbalanced moments of the inner gimbal; TOm is the inertial torque induced by unbalanced moments of the outer gimbal; TIx, TIz are the reaction torques exerted on the inner gimbal by the outer gimbal; TOy, TOz are the reaction torques exerted on the outer gimbal by the platform; TP, TY are the control torques for pitch and yaw axes, respectively; the total friction and spring torques TIfc, TOfc exerted, respectively, about yaw and pitch axes, are assumed to be described as: Yaw-pitch gimballed seeker S" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003798_icem49940.2020.9270715-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003798_icem49940.2020.9270715-Figure2-1.png", + "caption": "Fig. 2. 3D-Winding structure of high-speed SRM, (a) Round copper wire, (b) Radial flat copper wire, (c) Tangential flat copper wire.", + "texts": [], + "surrounding_texts": [ + "In this section, the effects of different winding structures on motor performance will be analyzed. Firstly, three winding structures will be compared and electromagnetic torque will be calculated. Secondly, the effects of different motor sizes on the winding structure will be studied furtherly, and the range of motor sizes applicable to different winding structures will be proposed. Finally, an optimization method for winding structure design will be proposed and verified by finite element method (FEM)." + ] + }, + { + "image_filename": "designv11_34_0000485_(asce)as.1943-5525.0000559-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000485_(asce)as.1943-5525.0000559-Figure1-1.png", + "caption": "Fig. 1. (Color) Definition of the coordinates", + "texts": [ + " It is clear that A is symmetric, and L is symmetric, positive, and semidefinite when the undirected graph is connected [see e.g., Ren et al. (2007) for more]. In this paper, the free-flying spacecraft is treated as a reference spacecraft. To describe the absolute position of the reference spacecraft, an ECI frame named J2000 is used, while to describe the relative motion between every two spacecraft, an orbital reference frame named the local vertical/local horizontal (LVLH) frame is used. As shown in Fig. 1, OEXIYIZI denotes the ECI frame with the origin OE attached to the mass center of the Earth; ORXRYRZR denotes the LVLH frame with the origin OR attached to the mass center of the reference spacecraft; Point Oi denotes the mass center of the ith maneuvering spacecraft; and vector \u03c1i denotes the relative position of the ith maneuvering spacecraft with respect to the LVLH frame. The LVLH frame is defined such that the XR axis is outward along the radial (local vertical), YR is perpendicular to XR in the orbit plane in the direction of motion (local horizontal), and ZR is along the orbit normal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure11-1.png", + "caption": "Fig. 11. Definitions of investigated points and surfaces of the wave cooling duct (left) and the spiral cooling duct (right). The figure shows a cross section of the ducts.", + "texts": [ + " 9 for the operating point (5000 rpm, 48 Nm). The air gap diameter is 160.4 mm (outer rotor) and 161.9 mm (inner stator). The speed 5000 rpm corresponds to a rotor peripheral speed of 5*80.2*\u03c0/60=21 m/s, so simulations agree with calculations, see Fig. 9c of air-gap velocity on the xy-plane (in the middle of the motor). The heat transfer coefficients of the wave and spiral cooling ducts are investigated for different flow rates, showing slightly higher values for the spiral duct, see Fig. 10. Temperatures in two points (see Fig. 11) and heat flux on different duct surfaces (see Fig. 12) are simulated from the wave and spiral cooling duct models in order to give input to calculations of thermal resistances for the LPN model. It can be seen in Fig. 12 that the ratio of heat transfer distribution between the inner and outer sides changes with flow rate. At higher flows, a larger portion of the heat goes through the inner duct surface relative to the outer duct surface. (a) Temperatures. (b) Water velocity in the cooling duct" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure10-1.png", + "caption": "Fig. 10. Rotation \u03be i and tangent vectors at the contact points.", + "texts": [ + " This can be done by imposing that the tangent vectors to the contact points (denoted by t R and t F ) and the normal to the ground n are orthogonal: n \u00b7 t R = 0 , n \u00b7 t F = 0 . (24) In order to obtain the position vectors of the contact points r R and r F , and the corresponding tangent vectors, nongeneralized coordinates \u03be R and \u03be F are defined. These coordinates represent a counterclockwise rotation about the Y 4 and Y 7 axes, orientating axes X \u03beR and X \u03beF so that they point to the contact points. The corresponding rotation matrix is: R \u03bei = \u239b \u239d cos (\u03bei ) 0 sin (\u03bei ) 0 1 0 \u2212 sin (\u03bei ) 0 cos (\u03bei ) \u239e \u23a0 , i = R, F . (25) Fig. 10 shows this rotation and the tangent vectors t R , t F . Therefore, r R and r F are given by r R = r G 4 + R 41 R \u03beR \u239b \u239d R 0 0 \u239e \u23a0 , r F = r G 7 + R 71 R \u03beF \u239b \u239d R 0 0 \u239e \u23a0 , (26) where R is the wheel radius; r G 4 and r G 7 are the absolute position vectors of the centers of mass of bodies 4 and 7; and R 41 , R 71 are the orientation matrices of body frames 4 and 7. With regard to t R and t F , they are computed as follows: t R = R 41 \u2202 \u0304r 4 G 4 R \u2202 \u03beR = R 41 \u239b \u239d \u2212R sin ( \u03beR ) 0 \u2212R cos ( \u03beR ) \u239e \u23a0 , t F = R 71 \u2202 \u0304r 7 G 7 F \u2202 \u03beF = R 71 \u239b \u239d \u2212R sin ( \u03beF ) 0 \u2212R cos ( \u03beF ) \u239e \u23a0 , (27) where r\u0304 4 G 4 R and r\u0304 7 G 7 F , expressed in body frames 4 and 7, are given by: r\u0304 4 G 4 R = \u239b \u239d R cos ( \u03beR ) 0 \u2212R sin ( \u03beR ) \u239e \u23a0 , r\u0304 7 G 7 F = \u239b \u239d R cos ( \u03beF ) 0 \u2212R sin ( \u03beF ) \u239e \u23a0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003437_s00170-020-05886-7-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003437_s00170-020-05886-7-Figure6-1.png", + "caption": "Fig. 6 Flexure mechanism for vertical motion of the microscope objective, a plan view, b orthographic projection, and c elevation showing the flexure hinge points as circles [34]", + "texts": [ + " [34] presented a new methodology to overcome several restrictions of 3D-printed mechanisms by exploiting the compliance of the plastic to produce a monolithic 3D-printed flexure translation stage. This structure is capable of sub-micron-scale motion over a range of 8 \u00d7 8 \u00d7 4 mm. Parallelogram structures form the basis of the microscope mechanism: the four-bar linkage enables the objective to vertically translate while not changing orientation or lateral position. This is close to the yield strain of both the PLA and ABS plastics which are typically utilized in 3D printing. They implemented a simple optical microscope based around the printed translation stage (Fig. 6) to quantify their mechanical performance in a realistic situation. This allows us to measure its stability over a range of time scales and to demonstrate the precision which can position a sample relative to the objective lens. Wei et al. [31] reported the fabrication and characterization of a flexure parallel mechanism through the laser beam melting (LBM) process to demonstrate the development of the proposed AM technique. The geometrical accuracy of the AM flexure mechanism was investigated by three-dimensional scanning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002272_2016-01-1958-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002272_2016-01-1958-Figure4-1.png", + "caption": "Figure 4. Geometric model and generated mesh of outer ring, bolt, dust shield", + "texts": [ + " The commercial software, CATIA [5], was used to design the outer ring, bolt, shield and knuckle. Because the hub and inner ring of the wheel bearing have no significant effect, they were not considered in this study. Due to high contact stresses between the rolling elements and raceways, the raceways are heat treated, and more specifically, induction hardened, a common process performed to improve strength and hardness selectively at the contact surfaces. The geometrical model and generated mesh are displayed in Figure 4. Figure 4(a) identifies the induction heat treated (hardened) area and the non-heat treated (base) area. These were modeled separately. Figure 4(b) shows the mesh which was generated using the commercial software, Hypermesh. [6] The total number of nodes and elements used were 103,000 and 255,000, respectively. The number of eight-node hexagonal elements used was 43,000. For the four-node tetrahedron elements, 212,000. and knuckle The outer ring was modeled using SAE1055 material properties. Different material properties were applied to the heat treated area in comparison to the non-heat treated area. Elasto-plastic analysis was performed to compute the distortion behavior of the outer ring, bolts and knuckle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003308_s40430-020-02488-y-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003308_s40430-020-02488-y-Figure5-1.png", + "caption": "Fig. 5 Principle of the cutting method", + "texts": [ + " This method is convenient for parameterized modeling and makes a variety of gear models to meet the requirement by modifying the parameters. Moreover, this method is considered to be applied in computers, so its application is very extensive. However, the sweeping method cannot solve the geometric relationship between the tooth profile and the sweep line. Therefore, a certain precision deviation exits between the mathematical model, established by this method, and the actual SBG tooth surface. The cutting principle method is based on the principle of gear cutting, as shown in Fig.\u00a05. Firstly, the shape of the surface of the hob and motion forms the imaginary generating gear. When the gear billet and the imaginary generating gear are 1 3 meshed, the reduction operation of the Boolean operation is carried out. After that, the gear billet is rotated to a certain angle, and the reduction operation of the Boolean operation is repeated. This series of steps forms a three-dimensional model of the gear. According to the global conjugate theory and the meshing equation, the imaginary generating gear is in line contact with the gear billet during the meshing process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001094_gt2014-25395-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001094_gt2014-25395-Figure10-1.png", + "caption": "FIGURE 10. Design concept of the test rig", + "texts": [ + " The results show that instable behaviour starts before reaching the critical frequency. The area of instability decreases with a higher damping ratio. Right after passing the critical frequencies a possible rub is stable. Beside the stability thresholds the computed p/q ratio is plotted. This ratio reflects the estimation of critical frequency-ranges, where a rub would lead to instable behaviour. To determine and validate the calculated parameters and the predicted rotor behaviour a test rig is scheduled. Figure 10 displays its main components. Since the main focus are rotordynamic considerations the test rig will be open with no fluid passing through the seal. Most importance is attached to the reflection of thermal conditions and vibration characteristics under the influence of eccentric rubbing brush seal. A major request is the feasibility to generate a rotating thermal bow on an elastic rotor under different rub conditions. The idea is to separate these requirements: rub induced heat generates a thermal bow on a thick rigid drum, which is supported by compliant shaft ends providing elasticity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure5.5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure5.5-1.png", + "caption": "Figure 5.5 Zero-offset version of the manipulator with an RRP arm.", + "texts": [ + " For the present manipulator, the sign variable \ud835\udf0e, which appears in the expressions of \ud835\udf0301 and \ud835\udf0312, represents two distinct optional inverse kinematic solutions that correspond to the same specified location of the wrist point R. The optional poses of the manipulator described by \ud835\udf0e = + 1 (left-shouldered pose) and \ud835\udf0e = \u2212 1 (right-shouldered pose) are sketched in Figure 5.4. Note that the availability of such optional poses becomes advantageous especially when there are obstacles to be avoided in the vicinity of the manipulator. (e) Position Singularity Consider a special version of the present manipulator such that d12 = 0. It is shown in Figure 5.5 in its side and front views. Regarding Eq. (5.93) for this special manipulator, it is possible to have r2 1 + r2 2 = 0 or r1 = r2 = 0. Such a particular pose of the manipulator is called a position singularity. In such a pose, Eq. (5.92) implies that the third link assumes a vertical pose so that s\ud835\udf0312 = 0, i.e. \ud835\udf0312 = 0 or \ud835\udf0312 = \ud835\udf0b. However, \ud835\udf0312 = 0 is the only physically possible value due to the shapes of the relevant links. On the other hand, Eq. (5.93) degenerates into 0 = 0 at the singularity and therefore \ud835\udf0301 becomes indefinite" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002574_detc2016-60019-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002574_detc2016-60019-Figure8-1.png", + "caption": "Fig. 8 CAD model and a 3D layered based virtual AM model", + "texts": [ + " 13 Superposition of CAD model on its virtual AM model The proposed CAV tool is being developed within SOLIDWORKS through its macro programming interface. Three typical examples were presented in Figs. 8-9, 10-11 & 12-13. Typical prototypes geometry used for this study is either composed of planer surfaces or free-form shapes and planer surfaces at different orientations. The contour data is generated in slicing stage through MATLAB. Then it is transformed into a macro file format for part quality analysis in a CAD environment. The example part 1 is shown in Fig. 8a and Fig. 8b in CAD model and layered model, respectively. The layer thickness of the 3D layered model is 0.254mm. The family of 2D contours is extruded into a 3D layered model in a CAD environment to detect the part quality issue for AM prototypes layer to layer. The part quality issued associated with this model are Shape Deviation (SD), Volumetric Error (VE), the Surface Roughness (Ra) and Feature Loss (FL). Each of the issues is identified individually through CAV tool for part quality analysis purpose. The layer model is shown in Fig. 8(b) with a higher shape deviation issues. The shape deviation of situations is automatically generated during resembling the pattern of AM process in a CAD environment as shown in Figs. 8b&9. For this example, layers with oversize, undersize, equal size and the combination of oversize and undersize situations are found over the virtual AM prototype surface. The nomenclature of the shape deviation situations are OSD, USD and MSD for the oversize, undersize and the combination of oversize and undersize situations, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000344_gt2015-43161-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000344_gt2015-43161-Figure10-1.png", + "caption": "Fig. 10 Experimental apparatus for measuring the threshold speed of instability of a rotor", + "texts": [ + " Figure 9 shows the effect of the damping coefficient of the flexible structure, bw, on the value of M* by varying the values of bw. As shown, the value of M* for larger values can be improved by increasing the damping coefficient of the flexible structure. 5 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use To confirm the validity of the numerical predictions, an experiment was conducted to measure the instability threshold of the proposed bearings. Figure 10 shows the experimental apparatus for measuring the threshold speed of instability. A rotor with a mass of 4.8 g was set vertically and was supported by aerostatic thrust bearings installed at the upper and lower ends of the rotor. Air jets injected from turbine nozzles drive the rotor. As seen, the bearing bush was flexibly supported by straight spring wires. The rotor displacement and the rotational speed were measured by two optical fiber sensors. In this experiment, we prepared two types of rotor with different bearing clearances, cr = 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002283_978-3-319-44156-6_1-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002283_978-3-319-44156-6_1-Figure5-1.png", + "caption": "Fig. 5 Finite element model: mesh and applied loads", + "texts": [ + " For the particular case studied in the present work, the results for the interferences can be seen in Fig. 4. As it has been mentioned before, the friction moment is an especially relevant magnitude in slewing bearings. Therefore, it would be interesting to evaluate the influence of the manufacturing errors in the friction moment. With this aim, different FE calculations were performed, which results are given and compared in the next section. The model used for the friction moment calculation is based on the one developed previously by the authors in [4], which is shown in Fig. 5. For the modelization, only the sector corresponding to one ball is considered. Note that the mesh was built putting special care to the contact region. Applied loads and boundary conditions are detailed below: \u2022 Outer ring: external surfaces fixed to the ground. \u2022 Ball preload: simulated by imposing a specific thermal condition. \u2022 Manufacturing tolerances: simulated by introducing an offset to the contacting surfaces of the raceway, as done by Aithal et al. [5]. This supposes an improvement over the previous model [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002591_s1068798x16120169-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002591_s1068798x16120169-Figure3-1.png", + "caption": "Fig. 3. Lateral gaps in the engagement of harmonic gears with \u03b1w = 20\u00b0 (1), 14\u00b0 (2), and 10\u00b0 (3).", + "texts": [ + " 2a) or outside it (Fig. 2b). In both cases, the azimuthal gap between the tooth profiles is , where is the central angle equal to the angular gap between involute curves In1 and Ina over the arbitrary circumference (radius rhy) of the hypothetical gear used in determining the azimuthal gap. As an illustration of the variation in the lateral gaps observed in the engagement of harmonic gear drives with Zf l = Z2 = 362, Zri = Z1 = 360, \u03b2 = 60\u00b0, x2 = 3, hc = 4.8 m, m = 0.5 mm we plot ji as a function of \u03d5 in Fig. 3. Vertical lines denote gaps at the onset of engagement. Analysis of Fig. 3 indicates that the position of the engagement zone of the f lexible gear\u2019s teeth relative to the deformation axis and the multiplicity of pairs in engagement depend on the geometric parameters of the harmonic engagement. The position of the zone where the f lexible gear rests on the harmonic generator may be determined on the basis of experimental data [4, 10, 11]. We may conclude that, for the given class of transmissions, this zone has fixed boundaries and only the overall reaction q(\u03d5) varies with increase in the load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000388_romoco.2015.7219734-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000388_romoco.2015.7219734-Figure11-1.png", + "caption": "Figure 11: FW-VTOL Tilt-rotor of VTOL Tech.", + "texts": [ + " One or several additional degree(s) of freedom allow one for the modification of the vehicle\u2019s shape with extended control possibilities. The Vertol VZ-2 Tilt-Wing of Fig. 1 is an early example of this class of vehicles with the main wing and propellers rigidly attached to each other but tilting with respect to (w.r.t.) the main body. This type of configuration can also be found at the MAV scale (like the SUAVI [4] or the QUX-02 [15] quad tilt-wings). Other types of tiltrotors/tilt-wings MAVs have been proposed, like the tilt-rotor FW-VTOL of VTOL Tech. (Fig. 11), or the tilt-wing ConvertISIR (Fig. 12), developed in our Lab. Concerning the FW-VTOL, the four propellers can be used in hover like for a classical quadrotor, and they tilt progressively pitch down for the transition to cruising flight. With respect to tilt-bodies aircraft, the main advantage is that the wing can always keep a small angle of attack, thus making the transition easier. It should be noted that this property is not satisfied for a convertible like the Vertol VZ-2. In this case, transition still requires large variations of the angle of attack despite the additional degree of freedom", + " By proceeding as in the previous cases, it is normally possible to simplify the expression of F and \u0393 by a change of control variables as follows (compare with (3)): F = F0(R, va)\u2212 TcRQ(s0)e3 + ns\u2211 j=1 Fj(R, \u03b4j , v a j ) \u0393 = \u03930(R, va) + \u0393c with Tc defined as above and \u0393c the new torque control variable. As in the case of tilt-body vehicles, some of the variables \u03b4j might be constrained, through the change of control variable, by the values of Tc and \u0393c. Some particular configurations have been more specifically investigated. Tilt-rotors: A first configuration is when the moving surfaces are only used for torque generation. An example is given by the FW-VTOL of Fig. 11. Then, as in the case of tilt-bodies, the term\u2211ns j=1 Fj(R, \u03b4j , v a j ) can essentially be neglected and the expression of F reduces to F = F0(R, va)\u2212 TcRQ(s0)e3 This expression is similar to the first equality in (3) with the additional control parameter s0. The important point is that the resulting force can be partially modified via s0 without modifying the main body\u2019s orientation (matrix R). This is the main advantage of tilt-body vehicles, which allows for transitions from hover to cruising flight while keeping a constant pitch angle of the main body" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.39-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.39-1.png", + "caption": "Figure 9.39 Second kind of motion singularity of the manipulator.", + "texts": [ + "621) If the task motion obeys the above velocity restriction, Eq. (9.605) leaves ?\u0307?1 free. However, owing to Eq. (9.609), this freedom allows ?\u0307?1 to be specified freely as desired. Thus, as a useful feature, this singularity helps the angular velocity of the end-effector to be fully specified as long as r\u20d7SP is kept parallel to the axis of the first joint. (b) Second Kind of Motion Singularity Equations (9.614) and (9.615) imply that the second kind of motion singularity occurs if cos\ud835\udf033 = 0 or \ud835\udf033 = \ud835\udf0e\u20323\ud835\udf0b\u22152 with \ud835\udf0e\u20323 = \u00b11. This singularity is illustrated in Figure 9.39 in its extended (\ud835\udf033 = \ud835\udf0b/2) and folded (\ud835\udf033 = \u2212\ud835\udf0b/2) versions. In this singularity, with \ud835\udf0323 = \ud835\udf032 + \ud835\udf0e\u20323\ud835\udf0b\u22152, Eqs. (9.612) and (9.613) can be manipulated into the following forms. (b2?\u0307?2 + \ud835\udf0e\u20323d4?\u0307?23)s\ud835\udf032 = \u2212v\u22171 (9.622) (b2?\u0307?2 + \ud835\udf0e\u20323d4?\u0307?23)c\ud835\udf032 = \u2212v\u22173 (9.623) The above equations become consistent and thus they can give finite values for ?\u0307?2 and ?\u0307?23 if the following compatibility condition is satisfied by the planned task motion. v\u22171c\ud835\udf032 = v\u22173s\ud835\udf032 (9.624) The same equations can be combined into the following equation, which relates the finite but indefinite values of " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001692_1.b35831-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001692_1.b35831-Figure1-1.png", + "caption": "Fig. 1 Representations of a) blade/casing rub in a turbomachinery stage, and b) metal cutting in a multiflute milling operation.", + "texts": [ + " We quote from a paper on the state of the art in turbomachinery coatings [15]: \u201cThe blade tips spinning at high speeds must act as an efficient cutting tool. This is necessary so as not to damage the seal or blades which can result in blade fracture if inefficient cutting is experienced.\u201d The context of a proper cutting operation can also be found in [14], where the authors state that blade tips are coated with abrasive cubic boron nitride particles to facilitate the cutting action. The similarity between blade/casing rub and internal milling dynamics is further illustrated in Fig. 1. Notice here that the red marked area indicates the material to be removed. Multiple blades rubbing against the coating resemble amultiflute milling cutter that is engaged in aworkpiece duringmachining. This apparent similarity is the theme in some recent studies [16,17]. The former is essentially a preceding study to this paper where the analogy and its repercussions on the stability are investigated. The latter also considers a similar approach but differs significantly in the aspects of dynamic modeling and stability assessment methodology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002262_978-3-319-27333-4-Figure3.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002262_978-3-319-27333-4-Figure3.1-1.png", + "caption": "Fig. 3.1 Bearing geometry with coordinate system", + "texts": [ + "11), the following boundary conditions are used: At Ly T T= =0, (3.11a) At Uy h T T= =, (3.11b) 3.2 Mathematical Analysis 22 and mT h T dy h = \u00f2 1 0 (3.11c) Thus, temperature profile expression (written in (3.11)) takes the following form: T T T T T y h T T T y h = - + -( )\u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 + + -( )\u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7L L U m L U m4 2 6 3 3 6 2 (3.12) where, TL, TU, and Tm represent temperatures of the lower bounding surface (journal), upper bounding surface (bearing), and mean temperature across the film, respectively (Fig. 3.1). - \u00b6 \u00b6 \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 = ( ) -( ) = k T x h T l y T x l s s s c s s a s , k T y k T y y oil upper bounding surface s s s s \u00b6 \u00b6 \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 = \u00b6 \u00b6 \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 =0 - \u00b6 \u00b6 \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 = ( ) -( ) = k T y h T x t T y t s s s c s s a s , 3 Performance Parameters 23 - \u00b6 \u00b6 \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 = ( ) -( ) = k T x h T y T x s s s c s s a s , 0 0 T T= U T T= L Substitution of \u2018u\u2019, \u2018w\u2019, and \u2018T\u2019 expressions ((3.9), (3.10), and (3.12)) into energy equation (3.8) and subsequently integrating the energy equation across the film thickness from the limit \u20180\u2019 to \u2018h\u2019 yield the following form of energy equation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001626_j.proeng.2015.12.125-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001626_j.proeng.2015.12.125-Figure5-1.png", + "caption": "Fig. 5. The conical wheel non-involute tooth profile.", + "texts": [], + "surrounding_texts": [ + "Solid models of gears with more complex geometry surfaces of the teeth, you can create a combination of the two above-described software systems. For example, involute bevel wheel (Fig. 4). Involute-bevel gear (IBG) - the wheel, sliced tool rack type (Gear comb, hob, grinding wheel) with variable displacement along the axis of the wheel tool [2]. A feature of such wheels is that in each face section profile is obtained with a certain offset coefficient which is changed in each section by the amount x = s tg / m, (1) where s \u2013 move; x \u2013 increment of bias; \u2013 taper angle IBG; m \u2013 module. Coefficient of radial clearance: (2) The coefficient of the tooth head height tool: cos cos** aat hh , (3) Socket module: cos m mt . (4) With the help of emulation tooth profile rack-type tool to create a number of profiles with a certain step s along the axis of the wheel. In Autodesk Inventor, create work planes at a distance from each other at a pitch s, and copy to the appropriate profiles. IBG get a solid model using the Loft applications." + ] + }, + { + "image_filename": "designv11_34_0003098_b978-0-12-821354-4.00009-1-Figure9.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003098_b978-0-12-821354-4.00009-1-Figure9.2-1.png", + "caption": "FIG. 9.2", + "texts": [ + " Reproduced with permission from Jin, H., Guo, C., Liu, X., Liu, J., Vasileff, A., Jiao, Y., Zheng, Y., Qiao, S.-Z., 2018. Emerging two-dimensional nanomaterials for electrocatalysis. Chem. Rev. 118, 6337\u20136408. Copyright (2018) American Chemical Society. properties of the resultant hybrid nanosystems are not only the result of the sole properties of any one component, but the degree of the interactions between the components considerably contributes to the final properties of the hybridized nanosystems. Fig. 9.2 shows the numerous optimization protocols of the hybridization of materials, which leads to modifying the physical, electronic, chemical, and surface properties that directly alter their inherent performance. Also, it makes it possible to introduce new physical properties and one or more novel functionalities that were absent earlier in any individual original component of hybrid nanosystems, for example, magnetic, electrical, optical, hydrophobicity, and hydrophilicity characteristics (Manias, 2007)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002074_978-3-642-36282-8-Figure4.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002074_978-3-642-36282-8-Figure4.4-1.png", + "caption": "Fig. 4.4: a) Flux distribution in an asynchronous machine during no load operation; b) flux distribution with locked rotor", + "texts": [ + " The interaction of non-sinusoidal inductions is represented by the interaction factor kinter. This factor is in the range of 10 %. ( ) \u22c5\u22c5+\u22c5=\u22c5+\u22c5== Fe 71.1 1 36.1 interFe,1interFe1Fe,Fe,F 11 k e kPkkPPP \u03bd \u03bd \u03bd \u03bd \u03c8 \u03c8\u03bd . (4.21) 4.3 Additional Iron Losses within the Stator and Rotor Lamination 51 The flux distribution for the fundamental and the higher harmonics are not the same. The fundamental consists of \u201cmain flux\u201d and \u201cleakage flux\u201d, the higher harmonics are damped by rotor currents. Mainly leakage flux exists, see fig. 4.4 b). The field distribution and the iron losses in stator and rotor have been calculated for the numerical field calculations fig. 4.4 a, b. The losses are given in table 4.3. 52 4 Additional Losses Due to Higher Voltage Harmonics The non-linear calculated losses during no load operation are about 15% smaller than for the locked rotor, if the voltage is kept the same. The additional losses are especially generated in the stator teeth. Additional losses do occur on any conducting design parts within electrical machines, which are exposed to magnetic flux. These parts can be surfaces of the rotor in synchronous machines, press-plates, iron housings, support rings or end bells" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002537_s11029-016-9608-x-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002537_s11029-016-9608-x-Figure1-1.png", + "caption": "Fig. 1. Schematics of a three-layer test specimen with an end part (a), a three-layer two-arm test specimen (b), and testing a two-arm test specimen on an electrodynamic vibration shaker (c).", + "texts": [ + " In accordance with this standard, the dynamic behavior of cantilevered test specimens of various structure is investigated by the acoustic resonance method. Without distorting the overall standard procedure, the authors of this paper made some modifications in the conditions of experimental studies and the instrumental and hardware tools used. First, as indicated in the standard, it is important that test specimens be fixed rigidly. Therefore, it is recommended to use specimens with a pronounced part end to be fixed in grips (Fig. 1a). Investigation results showed that the fixation significantly affects the natural frequencies and damping characteristics of test specimens. This effect is especially observed in dynamic tests of specimens with a soft layer. Therefore, we propose a variant of the experimental setup where the rigid fixation is realized more strictly by using a two-arm test specimen whose middle part is located between two calibrated cylinders (Fig. 1b) Second, the experimental study of the dynamic behavior of test specimens were carried out on an electrodynamic vibration setup (Fig. 1c) ensuring a highly precise specified vibration amplitude A0 of exciter and adjustment of its frequency f according to the resonance for the lowest vibration mode of test specimen. The kinematic response A of specimen end to forced vibrations was recorded by a RIFTEK firm (RF603-X/100) laser sensor ensuring measurements of vibration amplitudes in a digital format accurate to 0.01 mm. The mathematical software developed allowed one to carry out up to 2000 measurements of deflection per second, which secured a highly accurate description of experimental vibration records of tests specimens", + " Determination of the elastic characteristics of soft materials in shear is reduced to a sequence of solutions of the direct problem, which consists in obtaining the calculated amplitude A of resonance vibrations at a free end of test specimen from the specified amplitude of exciter and known elastic and damping properties of the material upon reaching the minimum of the assigned difference between the calculated and experimental vibrations amplitudes. To solve the direct problem, the finite-element method with three-layer beam elements for the test specimen is used. It is assumed that the outer layers of the element deform within the framework of Kirchhoff\u2212Love hypotheses, while the inner, soft layer operates in transverse shear. The test specimen is connected with a system of Oxz coordinates (see Fig. 1b) moving translationally together with the exciter according to the law w t A ei t 0 0( ) = \u03c9 , where \u03c9 \u03c0= 2 f is the angular frequency of exciter coinciding with the lowest natural frequency of the test specimen. The nodal displacements of a finite element (see Fig. 2) with respect to this system of coordinates are determined by the vector { } { }( )r w we = 1 1 1 2 2 2\u03d5 \u03b8 \u03d5 \u03b8 , where w1 and w2 are deflections of nodes 1 and 2; \u03d51 and \u03d52 are rotation angles of the normal; \u03b81 and \u03b82 are rotation angles of the soft layer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.14-1.png", + "caption": "Figure 9.14 Left and right shouldered poses of a Stanford manipulator.", + "texts": [ + " (a) First Kind of Multiplicity The first kind of multiplicity is associated with the sign variable \ud835\udf0e1 that arises in the process of finding \ud835\udf031 by using Eq. (9.142). The manipulator attains the same location of the wrist point by assuming one of the two alternative poses that correspond to \ud835\udf0e1 = + 1 and \ud835\udf0e1 = \u2212 1. These poses are very similar to those illustrated in Figure 9.3 for the Puma manipulator. Therefore, they are also designated as left shouldered if \ud835\udf0e1 = + 1 and right shouldered if \ud835\udf0e1 = \u2212 1. They are illustrated in Figure 9.14. (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the wrist joint variables (\ud835\udf035, \ud835\udf034, \ud835\udf036). Therefore, it is the same as the third kind of multiplicity of the Puma manipulator. In other words, as illustrated in Figure 9.5, it occurs similarly as a wrist flip phenomenon without any visual distinction. A Stanford manipulator with d2 > 0 may have only one kind of position singularity. Equation (9.149) implies that it is the same as the third kind of position singularity of the Puma manipulator, which is explained in Section 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003043_physreve.101.043001-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003043_physreve.101.043001-Figure7-1.png", + "caption": "FIG. 7. Deformed shape of a disk with N = 3 creases and n = 1 hinge lines per half-face. A fixed angle, \u03b2, is imposed across each crease line. The facets separate by an angle \u03c6 and the rotation angle between facet i and i + 1 is \u03b8i. The center of the disk is moved down/upward with all creases forming an angle to the x-y plane. A side view is shown in (a) and (b) shows a view of along the normal of facet 2.", + "texts": [ + " We number the facets clockwise, with facet i located between hinges i \u2212 1 and i. The centerline of facet n + 1 is aligned with the line of symmetry. A Cartesian coordinate system is located at the center of the undeformed disk with the y axis aligned radially along a crease and the z axis pointing out of the page. Initially, the facets on each side of a crease are rotated about the crease line by an angle of \u03b2/2 from the x-y plane. This leaves an angle \u03b2 across each crease line, with the apex oriented in the positive z direction (i.e., mountain folds), as shown in Fig. 7. We assume the radius, R, is sufficiently large such that the crease angle, \u03b2, is effectively fixed, forming a boundary condition on the deformation of the adjoining faces [13]. The facets must separate and move from the planar state to accommodate the imposed crease angle. We allow the hinges to both rotate and stretch and impose these hinge deformations as boundary conditions on the subsequent facet, then derive the energy cost of the corresponding continuum shell deformations. The equilibrium shape is thus obtained by minimizing the total strain energy with respect to the hinge rotations and stretch angle. The analysis proceeds by incrementally moving the center of the disk down/upward, causing the creases to form an angle relative to the x-y plane, as shown in Fig. 7. At each deformation increment, the equilibrium shape and strain energy are obtained. We first consider the kinematics of this deformation before deriving the energy cost. 1. Kinematics The kinematics are described by writing the position of facet i + 1 relative to facet i, forming a kinematic chain. We label the unit normal vector of facet i: \u03b7i; and unit vectors along the facet edges: pi,1 and pi,2, as shown in Fig. 7. The normal and edge vectors of the first facet, adjacent to the crease, are set by the crease angle, \u03b2, the facet angle, \u03b1, and the imposed deformation angle, : \u03b71 = \u23a1 \u23a3 sin \u03b2 2 \u2212 cos \u03b2 2 sin cos \u03b2 2 cos \u23a4 \u23a6 p1,1 = \u23a1 \u23a3 0 cos sin \u23a4 \u23a6 p1,2 = \u23a1 \u23a3 sin \u03b1 cos \u03b2 2 sin \u03b1 sin \u03b2 2 sin + cos \u03b1 cos cos \u03b1 sin \u2212 sin \u03b1 sin \u03b2 2 cos \u23a4 \u23a6. (1) Facet i + 1 is initially aligned in the same plane as facet i, then rotated by an angle \u03b8i about the vector pi,2. Facet i + 1 is then rotated about its normal vector (\u03b7i+1) by an angle \u03c6, as shown in Fig. 7. The stretch angle, \u03c6, is assumed to be constant for all facets, which implies constant circumferential strain. Making use of Rodrigues\u2019 rotation formula, the orientation of facet i + 1 is expressed relative to facet i as \u03b7i+1 = \u03b7i cos \u03b8i + (pi,2 \u00d7 \u03b7i ) sin \u03b8i + pi,2(pi,2 \u00b7 \u03b7i )(1 \u2212 cos \u03b8i ), pi+1,1 = p1,2 cos \u03c6 + (\u03b7i+1 \u00d7 p1,2) sin \u03c6 + \u03b7i+1(\u03b7i+1 \u00b7 p1,2)(1 \u2212 cos \u03c6), pi+1,2 = pi+1,1 cos \u03b1 \u2212 (\u03b7i+1 \u00d7 pi+1,1) sin \u03b1 + \u03b7i+1(\u03b7i+1 \u00b7 pi+1,1)(1 \u2212 cos \u03b1). (2) 043001-4 The deformation of each face is symmetrical about its bisector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002627_icmic.2016.7804248-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002627_icmic.2016.7804248-Figure4-1.png", + "caption": "Fig. 4. Sensors location : Inclinometers A1, A2, A3 and home-made displacement sensors B1, B2, B3 and B3\u2019. Inset : home-made displacement sensor based on double optical encoder containing two CNY70 components.", + "texts": [ + " 1T H1 LD 2 T 2 L 2 D1 SA XX Ytg LD2 LLDcos90 (8) Hip flexion-extension angle \u03b1H between thigh and pelvis segments is calculated using equation (9) TD 2 L 2 D 2 T1 1T H1 P 211 H LD2 LDLcos XX Ytan L XXsin90 (9) Body segments lengths used in previous formulas have been deduced from subject\u2019s body height using standard human anthropometric data presented in table I. C. Instrumentation The angle calculation method proposed in this work is based on the varying distance measure between the extremities of two adjacent human body segments, as for example, the distance between shoulder and wrist. For this purpose, homemade displacement sensors noted Bi (i=1 to 3) as illustrated in Fig. 4 have been developed. Each sensor Bi (i=1 to 3) was based on double optical encoder containing two CNY70 reflective optical sensors with transistor outputs (See inset in Fig. 4). B1, B2 were positioned at the arm upper extremity to measure the displacement between shoulder and wrist of left (DUL) and right (DUR) upper limbs. B3 was placed between the rear of the rowing machine and pelvic segment to measure horizontal displacement X2 of the thorax-pelvic joint. B3\u2019 was placed on the rear of the machine to measure the seat displacement X1. To measure forearms and thorax inclinations, three GY-85 IMU have been used (A1, A2, and A3 in Fig.4.). Each IMU was composed of an ITG3205 triple-axis MEMS gyroscope, an ADXL345 MEMS triple-axis accelerometer [18], and a HMC5883L MEMS triple-axis digital compass. All of these MEMS use the I2C communication protocol. A1 and A2 were used to measure inclinations FL and FR for the left and right forearms respectively. A3 measured thorax inclination T in sagittal plane. To reduce noise and improve sensors responses, complementary filters [19] were applied for data fusion between accelerometer and gyroscope signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000318_j.finel.2015.07.008-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000318_j.finel.2015.07.008-Figure7-1.png", + "caption": "Fig. 7. FE-model of the turbo-chiller rotor-bearing system in Ref. [16].", + "texts": [ + " (21) and (22) can be rewritten as Fy Fz ( ) \u00bc Fy0 Fz0 ( ) \u00fe2B\u03b7\u03a9 \u03c83 min \u03b3nyy 0 0 \u03b3nzz \" # y z \u00fe 2B\u03b7 \u03c83 min \u03b2n yy 0 0 \u03b2n zz 2 4 3 5 _y _z \u00f025\u00de where the stiffness and damping coefficients of the bearing can be described as cyy czz ( ) \u00bc 2B\u03b7\u03a9 \u03c83 min \u03b3nyy \u03b3nzz ( ) \u00f026\u00de kyy kzz ( ) \u00bc 2B\u03b7 \u03c83 min \u03b2n yy \u03b2n zz 8< : 9= ; \u00f027\u00de The governing equation of the whole geared rotor system is established at last as follows, in which each node has 6 degrees of freedom, M \u20acX\u00fe\u00f0C\u00fe\u03a9G\u00de _X\u00feKX \u00bc F \u00f028\u00de where M is the mass matrix; C is the damping matrix; G is the gyroscopic matrix; \u03a9 is the spin speed of rotor; K is the stiffness matrix; F is the force vector including unbalanced force vector and external additional loads; X is the displacement vector of each node, which can be written as Eq. (29), and n is the number of nodes. X \u00bc \u00bdu1; v1;w1;\u03b8x1;\u03b8y1;\u03b8z1;u2; v2;w2;\u03b8x2;\u03b8y2;\u03b8z2;\u2026;un; vn;wn;\u03b8xn;\u03b8yn;\u03b8zn T \u00f029\u00de With the proposed 6-DOF modeling method, an unbalance response analysis is performed with a 600 me turbo-chiller rotor bearing system in Ref. [16], which is shown in Fig. 7. When a test unbalance of 19.7 a mm is attached to the impeller, the maximum coupled unbalance responses at the motor rotor through the 6-DOFs modeling method in this paper together with the results of Ref. [16] are shown in Fig. 8. The unbalance responses in this paper are consistent with the results in the reference. The FE model of the five-shaft geared rotor system is shown in Fig. 9, where 169 beam elements and 174 nodes are included and shown in Table 1. The gears, couplings, and impellers are simplified as lumped mass elements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000999_0954406212473242-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000999_0954406212473242-Figure1-1.png", + "caption": "Figure 1. A calculation model for bioliquid-filled microtubule considering surface effect, where Es, , hs represent the surface modulus, residual surface tension, and thickness of surface layers of the bioliquid-filled microtubule, respectively.", + "texts": [ + " In examples, the influence of surface modulus, residual surface tension and bioliquid density on the coupling frequency behavior of bioliquid-filled MT are described and discussed. The new study results may be a useful reference for the ultrasonic examine of nano-MT organization with surface effect, and may be applied in some biomedical clinical fields. Coupling vibration equations considering surface effects The structure of MT is composed of laterally parallel protoElaments. As the inner radius and the outer radius of MT are very close, a bioliquid-filled MT may be described as a bioliquid-filled hollow beam with consideration of surface effects, which is shown in Figure 1. In general, the length of MT is much larger than its outer radius, so that the shearing deformation is not considered in the beam model.14 Surface energy in nanomaterial is composed of two additional distinct effects,22 which are, respectively, the surface modulus and surface residual tension in two surface layers of MTs, as shown in Figure 1. The surface modulus in two surface layers with nearly zero thickness may be idealized to an equal constant value Es (N/m) as a material property.32 The residual surface tension (N/m) will induce a transverse loading q\u00f0x\u00de applied on MT along the length direction.29 The equilibrium equation of bioliquid-filled MT considering surface effects is written as @2M\u00f0x, t\u00de @x2 \u00fe A @2w\u00f0x, t\u00de @t2 \u00bc pf \u00f0x, t\u00de \u00fe q\u00f0x\u00de \u00f01\u00de where M\u00f0x, t\u00de is the bending moment exerted on the bioliquid-filled MT, w\u00f0x, t\u00de is the bending deflection of the bioliquid-filled MT, pf \u00f0x, t\u00de is the dynamic pressure (per unit length) exerted on MT induced by bioliquid in the MTs, and q\u00f0x\u00de is a transverse loading exerted on MTs induced by the surface residual stress of the MT. The positive direction of pf \u00f0x, t\u00de and q\u00f0x\u00de in equation (1) is along y-axis direction as at UNIVERSITY OF TOLEDO LIBRARIES on March 12, 2015pic.sagepub.comDownloaded from seen in Figure 1. In the above formula, and A are the density and the cross area of bioliquid-filled MT, respectively. Based on the bending theory of beam, the bending moment equation of bioliquid-filled MT is expressed as M\u00f0x, t\u00de \u00bc EeIew 00\u00f0x, t\u00de \u00f02\u00de where Ee and Ie are, respectively, the effective Young\u2019s modulus and inertia moment of MT considering surface effects, which are as follows EeIe \u00bc E s\u00f0R3 0 R3\u00de \u00fe 0:25 E\u00f0R4 0 R4\u00de \u00f03\u00de where R and Ro are, respectively, the inner and outer radii of MT, Es is the surface modulus of MT and E is Young\u2019s modulus of MT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002834_tdmr.2020.2966043-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002834_tdmr.2020.2966043-Figure2-1.png", + "caption": "Fig. 2. The 3D-printed polyamide scanner and the experimental setup, composed of a laser, coil driven scanner, target sheet, and a CMOS camera.", + "texts": [ + " Based on (2) through (3), we find the fundamental resonance to be f0 = 98 Hz. The maximum force applied on the scanner, corresponding to a total optical scan angle of 80 \u25e6 degrees, can be deduced based on the well-known Hooke\u2019s Law. Yet at resonance, a given force results in a displacement that is Q (quality factor) folds higher than the static case [18] Thus using i) Hooke\u2019s Law, ii) Q times (measured to be \u223c 40) magnified displacement at resonance and iii) Eq. (3), we calculate the maximum applied force to be 3.4 mN. Fig. 2 illustrates the experimental setup that was utilized in the cyclic testing of the scanners. A laser beam was directed onto the scanner and deflected toward the millimetric paper. The scanner was driven via the external coil at its fundamental resonance, forming a vertical line whose length was observed by a CMOS camera and biconvex lens pair. The scanners were driven at the TOSA values of 40\u25e6, 50\u25e6, 60\u25e6, 70\u25e6, and 80\u25e6 (3 scanners for each TOSA, a total of N = 15 scanners were tested). The variations in the scan-line, as well as the variations in the fundamental resonance, were periodically monitored until fracture or 108 cycles (\u224810 days) if the device continued functioning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001644_ijrapidm.2015.073548-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001644_ijrapidm.2015.073548-Figure6-1.png", + "caption": "Figure 6 Surface comparison CGAM: sample 2", + "texts": [], + "surrounding_texts": [ + "Table 3 shows the overall linear and angular dimensions of the produced sample. The figure of standard uncertainty is also presented per dimension and per machine. Consistent with the results displayed in Table 3, all the dimensional features in the Z-axis (i.e. dimensions related with the bottom plane of the produced parts) have shown a much higher deviation. In this regard, the average error value of dimension 6 in the CGAM produced parts has grown to 2.245 mm and is not displayed in the figure. Figure 4 shows the average linear dimensional error of the parts after the hybrid manufacturing process. The errors have been calculated by computing the difference between the average dimensions of the produced sample and the nominal value. Results show that the dimensional errors related with the bottom plane of the part (i.e. attaching plane on the additive machine) suffer from much higher deviations due to the warping effect typically present in material extrusion processes. The results displayed in Figure 4 and Table 3 demonstrate that linear dimensions and angular dimensions on the horizontal plane (XY plane) have very low deviations and the performance is similar in both machines. According to the general tolerance standard (ISO, 2000), the tolerance class designation of the PGAM sample has fallen under the category of \u2018fine\u2019. In the case of the CGAM sample, average results fall under the same tolerance category. However, standard uncertainty is higher and certain dimensions, such as numbers 9 and 11, show poor results falling under the category of \u2018medium\u2019. Figures 5 and 6 show the CAD to part surface comparison of PGAM sample 3 and CGAM sample 2, respectively. Red areas represent the dimensions which are above the nominal CAD model and blue areas represent dimensions which are below the nominal CAD model. Deviation to nominal CAD model is in the range of \u20130.40 mm to 0.40 mm for PGAM produced parts and \u20132 mm to 2 mm for CGAM produced parts. At the same time, the bottom view of the figures shows that warping is much more severe in CGAM produced parts, thus having a deviation maximum up to 3.4 mm. Consistently, with the trends previously displayed in Figure 4 and Table 3, the top view of the visualised samples demonstrates that dimensional deviations are much lower in all the post-processed surfaces (i.e. XY plane), hence subtractive processes are increasing the dimensional stability. Figure 7 shows the General Dimensions and Tolerances (GD&T) studied during this experiment, and the data displayed correspond to sample 3, manufactured with CGAM equipment. Average DG&T values and corrected standard deviations for small samples are shown in Table 4. The cylindricity and roundness tolerances in holes 1 and 2 (GD&T 1, 2, 3 and 4) show equivalent results for CGAM and PGAM parts. The top plane and front plane perpendicularity (GD&T 5), the front plane and middle plane parallelism (GD&T 6) and the top plane flatness (GD&T 7) show better results for PGAM parts. The majority of these GD&T features fall under the tolerance class K, according to standards, not depending on the type of machine used. Table 4 also shows that thermal warping had a negative effect on GD&T features, such as the bottom plane and top plane parallelism (GD&T 8), the bottom plane and front plane perpendicularity (GD&T 9) and the bottom plane flatness (GD&T 10). They have been computed using the best fit plane of the bottom plane of the produced sample, and therefore its deviation is much higher. The qualitative evaluation of the tested additive machines showed that the impact of initial layers over the overall geometry is fundamental when using CGAM equipment. Figure 8 shows how the raft of the CGAM system has been detached from the base plane after one hour of manufacturing due to the thermal warping, which creates stress that causes the detachment of the geometry from the base plate. The CGAM system failed several times during this experiment, and the manufacturing process was unpredictable. In addition, as it is shown in Figure 9, the filament of the CGAM produced part had bad cohesion for the subtractive process. All the CGAM produced parts showed delamination of the filament in contact with the milling cutter, which also affected the final geometry. Nevertheless, dimensional stability was substantially improved and, for instance, the stair-step effect was eliminated after the milling process." + ] + }, + { + "image_filename": "designv11_34_0000207_9781119454816.ch8-Figure8.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000207_9781119454816.ch8-Figure8.3-1.png", + "caption": "Figure 8.3 Two-bar truss under load.", + "texts": [ + "-lb of frictional moment in bearing is equal to 0.0025 rad of angle of twist, which, in turn, is equivalent to 1 \u2218F rise in temperature, the constants a, b, and c can be determined. By using Eq. (E7), the optimization problem can be stated as Minimize f (R, L) = 0.44R3L\u22122 + 10R\u22121 + 0.592RL\u22123 (E12) subject to 8.63R\u22121L3 \u2264 1 (E13) The solution of this one-degree-of-difficulty problem can be found as R* = 1.29, L* = 0.53, and f* = 16.2. Example 8.13 Design of a Two-Bar Truss The two-bar truss shown in Figure 8.3 is subjected to a vertical load 2P and is to be designed for minimum weight [8.33]. The members have a tubular section with mean diameter d and wall thickness t and the maximum permissible stress in each member (\ud835\udf0e0) is equal to 60 000 psi. Determine the values of h and d using geometric programming for the following data: P = 33,000 lb, t = 0.1 in., b = 30 in., \ud835\udf0e0 = 60 000 psi, and \ud835\udf0c (density) = 0.3 lb/in3. SOLUTION The objective function is given by f (d, h) = 2\ud835\udf0c\ud835\udf0bdt \u221a b2 + h2 = 2(0.3)\ud835\udf0bd(0.1) \u221a 900 + h2 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000329_b978-0-12-397945-2.00009-3-Figure9.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000329_b978-0-12-397945-2.00009-3-Figure9.2-1.png", + "caption": "Figure 9.2 Basic wave parameters in beating flagella. l, wavelength; y, bend angle; s, shear angle at the point S. P, principal bend; R, reverse bend. One beat cycle of flagellar oscillation consists of a pair of bends (P and R bends) (A, B). Bend angle and shear angle (B, C) are parameters showing microtubule sliding in beating flagella.", + "texts": [ + " The flash frequencies should be set slightly lower than the flagellar beating frequencies: at such flash frequencies, bending waves can be observed as if they are slowly propagating from base to tip along the flagellum (Fig. 9.1 from left to right images). For observation of the slowly beating flagella, reactivated at low ATP concentrations (less than 10 mM), a halogen lamp is used. 2. From the recorded images, wave parameters, such as wavelength (l), bend angle (y), and shear angle (s), shown in Fig. 9.2, are obtained. The wavelength is defined as a length consisting of a pair of bends (Fig. 9.2A): in the flagellum of sea urchin sperm, asymmetrical bending waves, consisting of a principal bend (P) and a reverse bend (R), the former larger than the latter, are alternately formed (Gibbons & Gibbons, 1972) (Fig. 9.2B). The bend angle is a parameter proportional to the net amount of microtubule sliding that has occurred at the bending region in question (Fig. 9.3), assuming that there is no twisting of the axoneme or longitudinal compliance of the microtubules (Brokaw, 1991). In the case of asymmetrically beating flagella, an average bend angle of the P and R bends are usually used as a useful parameter. For further analysis of the microtubule sliding in beating flagellum, the shear angle curves showing microtubule sliding as a function of position along the length of a flagellum are useful; they are obtained as a difference in angular orientation between the locus on the flagellum and the basal end of the flagellum (or sperm head), where there is assumed to be no sliding (Gibbons, 1981; Satir, 1968; Warner & Satir, 1974) (Fig. 9.2C). In Fig. 9.4, the left graph is an example of a shear curve showing the time-dependent changes of shear angles along the flagellum for a cycle of beating and the right diagram shows calculated wave forms reconstructed from the shear curve. Changes of shear angle between two sequential lines at a given position along the flagellum represent the sliding velocity (Brokaw, 1991). Analysis of the sliding velocity, as well as the amount of microtubule sliding, during flagellar beating is important to understand the regulation of dynein activity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure26-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure26-1.png", + "caption": "Fig. 26. Rotations \u03be i , \u03b7i and local frames.", + "texts": [], + "surrounding_texts": [ + "In order to define the toroidal surface, two angular coordinates are required, as opposed to the ring case, where only the non-generalized coordinate \u03be i was necessary. Therefore, the angular coordinates \u03be and \u03b7 for toroidal and poloidal directions are defined. Fig. 25 shows the geometry of the torus, with b being the major radius (distance from the center G of the torus to the center of the tube P ), and a the radius of the tube. The first step to construct the model is locating the centers of the bodies 4 and 7 (denoted by G 4 and G 7 ), whose absolute position vectors can be found in Appendix A. Starting from body frames 4 and 7 (for the rear and forward wheel, respectively), one can define a counterclockwise rotation \u03be i about the Y 4 , Y 7 axes in order to orientate axes X \u03beR and X \u03beF , so that they point to the center of the tube P . The corresponding rotation matrix is: R \u03bei = \u239b \u239d cos (\u03bei ) 0 sin (\u03bei ) 0 1 0 \u2212 sin (\u03bei ) 0 cos (\u03bei ) \u239e \u23a0 , i = R, F . (57) In contrast to the ring case, axes X \u03beR and X \u03beF do not point to the contact points R and F . Therefore, next issue is to locate these contact points departing from the center of the tube P . Placing axes X \u03beR and X \u03beF on this point, one can reach points R and F by making a clockwise rotation \u03b7i around axis Z \u03be so that X \u03b7i points to R, F . The corresponding rotation matrix R \u03b7i is: i R \u03b7i = \u239b \u239d cos (\u03b7i ) sin (\u03b7i ) 0 \u2212 sin (\u03b7i ) cos (\u03b7i ) 0 0 0 1 \u239e \u23a0 , i = R, F . (58) These rotations and the local frames are shown in Figs. 26 and 27 . The absolute position vectors of the centers of the tubes, denoted by r P R and r P F , and the absolute position vectors of the contact points, r R and r F , are shown in Fig. 28 and can be computed from: r P R = r G 4 + r G 4 P R = r G 4 + R 41 R \u03beR \u239b \u239d b 0 0 \u239e \u23a0 , r P F = r G 7 + r G 7 P F = r G 7 + R 71 R \u03beF \u239b \u239d b 0 0 \u239e \u23a0 , r R = r P R + r P R P R = r P R + R 41 R \u03beR R \u03b7R \u239b \u239d a 0 0 \u239e \u23a0 , r F = r P F + r P F P F = r P F + R 71 R \u03beF R \u03b7F \u239b \u239d a 0 0 \u239e \u23a0 . (59) The next step is to obtain the holonomic and non-holonomic constraints. The addition of the two non-generalized coor- dinates \u03b7R and \u03b7F means that the multibody model is defined by 15 coordinates: q = ( x 2 y 2 z 2 \u03b12 \u03b22 \u03b32 \u03b852 \u03b832 \u03b865 \u03b843 \u03b876 \u03beR \u03beF \u03b7R \u03b7F )T . (60) Since the number of coordinates has been enlarged, the number of constraints should also be increased so that the model remains with 5 degrees of freedom. On the one hand, as in the case of the multibody model with ring wheels, holonomic constraints arise from forcing the contact of the wheels with the ground, being the corresponding set: C h ( q ) = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d r R Z r F Z n \u00b7 t L R n \u00b7 t T R n \u00b7 t L F n \u00b7 t T F \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239c \u239c \u239c \u239c \u239d 0 0 0 0 0 0 \u239e \u239f \u239f \u239f \u239f \u239f \u239f \u23a0 , (61) where t L R , t L F are the longitudinal tangent vectors to the contact points, and t T R , t T F are the transversal tangent vectors, depicted in Fig. 28 . These are given by Eqs. (62) and (63) : t L R = R 41 \u2202 \u0304r 4 G 4 R \u2202 \u03beR , t T R = R 41 \u2202 \u0304r 4 G 4 R \u2202 \u03b7R , (62) t L F = R 71 \u2202 \u0304r 7 G 7 F \u2202 \u03beF , t T F = R 71 \u2202 \u0304r 7 G 7 F \u2202 \u03b7F , (63) where r\u0304 4 G 4 R and r\u0304 7 G 7 F are: r\u0304 4 G 4 R = R \u03beR \u239b \u239d \u239b \u239d b 0 0 \u239e \u23a0 + R \u03b7R \u239b \u239d a 0 0 \u239e \u23a0 \u239e \u23a0 = \u239b \u239d (b + a cos (\u03b7R )) cos (\u03beR ) \u2212a sin (\u03b7R ) \u2212(b + a cos (\u03b7R )) sin (\u03beR ) \u239e \u23a0 , (64) r\u0304 7 G 7 F = R \u03beF \u239b \u239d \u239b \u239d b 0 0 \u239e \u23a0 + R \u03b7F \u239b \u239d a 0 0 \u239e \u23a0 \u239e \u23a0 = \u239b \u239d (b + a cos (\u03b7F )) cos (\u03beF ) \u2212a sin (\u03b7F ) \u2212(b + a cos (\u03b7F )) sin (\u03beF ) \u239e \u23a0 . (65) On the other hand, the non-holonomic constraints appear again since the wheels are assumed to roll without slipping. As in the ring wheels model, velocities of rear and forward contact points v R , v F are computed with: { v R = v G 4 + R 41 ( \u03c9\u0304 41 \u00d7 r\u0304 4 G 4 R ) , v F = v G 7 + R 71 ( \u03c9\u0304 71 \u00d7 r\u0304 7 G 7 F ) . (66) Therefore, the set of non-holonomic constraints is given by: C nh ( q , \u02d9 q ) = \u239b \u239c \u239c \u239d v R x v R y v F x v F y \u239e \u239f \u239f \u23a0 = \u239b \u239c \u239c \u239d 0 0 0 0 \u239e \u239f \u239f \u23a0 . (67)" + ] + }, + { + "image_filename": "designv11_34_0002095_s12206-015-1145-3-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002095_s12206-015-1145-3-Figure7-1.png", + "caption": "Fig. 7. (a) Case a: configuration of a positive position error without an orientation error; (b) case b: configuration of a positive orientation error with a positive position error; (c) case c: configuration of a negative orientation error with a positive position error with a top contact; (d) case d: configuration of a negative orientation error with a positive position error with a side contact.", + "texts": [ + " Therefore, that assumption is must necessary, and the amplitude and speed of the perturbation that is used for MPCM is experimentally determined. A total of four contact states can occur during the performance of assembly task. Even cases where the axes correspond to each other and differ from each other are considered to be identical. Performance index R and constraints of contact force condition or maximum error are determined when input is orientation and position perturbation. Eqs. (2) and (3) show the process for calculating perform- ance index R under condition of case a in Fig. 7(a) ( )\u03b5 sin \u03c9t\u03c6\u03c6 = If \u03c0 0 t 2\u03c9 < < ( )z z z 1F F K r sin \u03c6= + ( )( ) ( )2 x x 2 y 1 2 z 1M M \u00b5Nr K r r 1 cos K r sin\u03c6 \u03c6= \u2212 \u2212 \u2212 \u2212 zN F= If \u03c0 3\u03c0 t \u03c9 2\u03c9 < < (2) ( ) ( )z z z 1 eF F K r y sin \u03c6= \u2212 \u2212 ( ) ( )( ) ( ) ( ) x x 2 y 1 e 2 2 z 1 e M M \u00b5Nr K r y r 1 cos K r y sin \u03c6 \u03c6 = + + \u2212 \u2212 \u2212 \u2212 ( )( ) ( )( ) ori \u03c0 3\u03c0 2\u03c9 2\u03c9 x z z x x z z x x\u03c0 0 \u03c9 R F F M M dt F F M M dt= \u2212 \u2212 + \u2212 \u2212\u222b \u222b (3) oriyp x 1 eK R \u03b1 y= = \u2212 ( )yy \u03b5 sin \u03c9t= ( )( ) ( )( ) \u03c0 3\u03c0 2\u03c9 2\u03c9 xo z z x x z z x x\u03c0 0 \u03c9 K F F M M dt F F M M dt= \u2212 \u2212 + \u2212 \u2212\u222b \u222b ( ) posxo y 4 eK R \u03b1 tan " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003363_j.mechmachtheory.2020.104027-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003363_j.mechmachtheory.2020.104027-Figure1-1.png", + "caption": "Fig. 1. Structural drawings of CVT chain.", + "texts": [ + " Then, this research applied static analysis to determine how the shape of the chain can reduce the vibration exciting force. Specifically, the research focused on the effects of the CVT chain pin cross sectional curve on CVT vibration and noise. The vibration and noise performance of two CVT chains with different pin cross sectional curves were predicted using a formulated geometric model [2] . The validity of this geometric model was then verified by measuring the noise of a prototype chain. Fig. 1 shows the structure of the CVT chain used in this study and Table 1 shows the main chain specifications and experimental conditions. A CVT chain is composed of multiple units. A single unit consists of pin A, pin B, and several links arranged in the z direction and fixed to pins A and B. In a chain-type CVT, the pins contact with the conical pulleys on the input/output shaft, and power is transmitted by clamping from the pulleys. Pin A of the k th unit and pin B of the k + 1 th unit are in rolling contact with each other on the cross-sectional curve of the pins. This enables the chain to bend. When the units are on a straight line as shown in Fig. 1 -(1), the contact point between pin A of the kth unit and pin B of the k + 1 th unit is defined as the origin O k of the kth unit. In this state, the forward direction of the unit is defined as the coordinates of the unit in the X k axis. When the unit bends ( Fig. 1 -(2)), the contact point C k between the pins moves S k +1 ( \u03b1k ) in the Y k +1 axis direction according to the bending angle \u03b1k . The bending characteristics of the chain used in this study change according to the crosssectional curve \u03b6 p 1 of pin A. This is because the cross-sectional curve of pin B is a straight line parallel to the Y k axis. A later section will explain how this characteristic affects the vibration of the chord. Fig. 2 is a schematic diagram showing the action of the chain as it enters a pulley" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001035_09205071.2014.963698-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001035_09205071.2014.963698-Figure3-1.png", + "caption": "Figure 3. The integral path W in the complex plane.", + "texts": [ + " 0 (16) Accordingly, for q0 > r1 \u00fe r2 (8) can be transformed to a complex integral in the t-plane: M \u00bc l0pr1r2 Z W cosh l \u00fe f1\u00f0 \u00det\u00bd cosh l f2\u00f0 \u00det\u00bd sinh 2lt\u00f0 \u00de J1 r1t\u00f0 \u00deJ1 r2t\u00f0 \u00deH 1\u00f0 \u00de 0 q0t\u00f0 \u00dedt (17) This process is similar to that of converting the real-axis integral of the vector potential into the complex integral, which was introduced by Sommerfeld [8,9]. By this procedure, the dependence of the real-axis integral on the origin can be removed, and consequently the integral path is able to be deformed in the complex t-plane. In (17), the integration path W starts from \u00fei1, encircling the pole \u00feip=2l anticlockwise and then moving towards \u00fei1 again (See Figure 3). Now that all the simple poles iak \u00bc kpi 2l ; k \u00bc 1; 2; 3; . . . are enclosed by W, (17) can be solved by residue theorem: M \u00bc l0pr1r2 l P1 k\u00bc1 cos ak f2 f1\u00f0 \u00de \u00fe 1\u00f0 \u00dek cos ak f2 \u00fe f1\u00f0 \u00de h i I1 r1ak\u00f0 \u00deI1 r2ak\u00f0 \u00deK0 q0ak\u00f0 \u00de; q0 r1 \u00fe r2 (18) In (18), the constraint between \u03b61 and \u03b62 is superfluous. When q0 r2 r1, letting f2 t\u00f0 \u00de \u00bc cosh l \u00fe f1\u00f0 \u00det\u00bd cosh l f2\u00f0 \u00det\u00bd sinh 2lt\u00f0 \u00de J1 r1t\u00f0 \u00deJ0 q0t\u00f0 \u00de (19) f2(t) is an even function of t. Similar to (15), we have Z e 1 f2 t\u00f0 \u00deH 1\u00f0 \u00de 1 r2t\u00f0 \u00dedt \u00fe Z1 e f2 t\u00f0 \u00deH 1\u00f0 \u00de 1 r2t\u00f0 \u00dedt \u00bc 2 Z1 e f2 t\u00f0 \u00deJ1 r2t\u00f0 \u00dedt (20) However, it should be noticed that the remainder of the integral does not vanish: lim e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002305_b978-0-12-405935-1.00010-1-Figure10.28-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002305_b978-0-12-405935-1.00010-1-Figure10.28-1.png", + "caption": "FIGURE 10.28 Curving flight of rotating spheres, in which F indicates the force exerted by the fluid: (a) negative Magnus effect; and (b) positive Magnus effect. A well-hit tennis ball with spin is likely to display the positive Magnus effect.", + "texts": [ + " An important point for the present discussion is that for supercritical Reynolds numbers the separation point slowly moves upstream, as evidenced by the increase of the drag coefficient after the sudden drop shown in Figure 10.24. With this background, we are now in a position to understand how a spinning ball generates a negative Magnus effect at Re < Recr and a positive Magnus effect at Re > Recr. For a clockwise rotation of the ball, the fluid velocity relative to the surface is larger on the lower side (Figure 10.28). For the lower Reynolds number case (Figure 10.28a), this causes a transition of the boundary layer on the lower side, while the boundary layer on the upper side remains laminar. The result is a delayed separation and lower pressure on the bottom surface, and a consequent downward force on the ball. The force here is opposite to that of the Magnus effect. The rough surface of a tennis ball lowers the critical Reynolds number, so that for a well-hit tennis ball the boundary layers on both sides of the ball have already undergone transition. Due to the higher relative velocity, the flow near the bottom has a higher Reynolds number, and is therefore farther along the Re-axis of Figure 10.24, in the range AB in which the separation point moves upstream with an increase of the Reynolds number. The separation therefore occurs earlier on the bottom side, resulting in a higher pressure there than on the top. This causes an upward lift force and a positive Magnus effect. Figure 10.28b shows that a tennis ball hit with underspin (backspin) generates an upward force; this overcomes a large fraction of the weight of the ball, resulting in a much flatter trajectory than that of a tennis ball hit with topspin. A slice serve, in which the ball is hit tangentially on the right-hand side, curves to the left due to the same effect. Presumably soccer and golf balls curve in the air due to similar dynamics. A baseball pitcher uses different kinds of deliveries, a typical Reynolds number being 1.5 105. One type of delivery is called a curveball, caused by sidespin imparted by the pitcher to bend away from the side of the throwing arm. A screwball has the opposite spin and oppositely curved trajectory, when thrown correctly. The dynamics are similar to that of a spinning tennis ball (Figure 10.28b). Figure 10.29 is a photograph of the flow over a spinning baseball, showing an asymmetric separation, a crowding together of streamlines at the bottom, and an upward deflection of the wake that corresponds to a downward force on the ball. The knuckleball, on the other hand, is released without any spin. In this case the path of the ball bends due to an asymmetric separation caused by the orientation of the seam, much like the cricket ball. However, the cricket ball is released with spin along the seam, which stabilizes the orientation and results in a predictable bending" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003584_s41315-020-00146-z-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003584_s41315-020-00146-z-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of Au\u2013Ni\u2013Pt micro-vehicle with singular off-center nanoengine. The micro-vehicle has a form of disk comprising of three different metals: Au, Ni, and Pt. Au- and Ni-disks are concentric. Pt forms a tubular cylinder, which is the nanoengine of the micro-vehicle and is only located at one side of the Au\u2013Ni disk", + "texts": [ + " The propulsion of the micro-vehicle is characterized in deionized (DI) water at three different hydrogen peroxide (H2O2) concentrations. Experimental observation presents the micro-vehicle moves forward circularly without any external control, which reveals that the micro-vehicle can realize the circular self-steering. 2 Circular self\u2011steering propulsion mechanism for\u00a0gold\u2013nickel\u2013platinum micro\u2011vehicle with\u00a0singular off\u2011center nanoengine An innovative micro-vehicle with singular off-center nozzle nanoengine is proposed. The schematic and the material composition of the micro-vehicle are illustrated in Fig.\u00a01. In detail, the micro-vehicle is composed of three different metals: non-catalytic Au, magnetic Ni, and catalytic Pt. Au and Ni are formed as a concentric disk with 12\u00a0\u00b5m in diameter. The thicknesses of Au- and Ni-disks are 0.2 and 0.1\u00a0\u00b5m, respectively. The singular off-center nozzle Pt nanoengine has a tubular shape with a solid base, which is located at one side of the Au\u2013Ni disk. The diameter, base-thickness, wallthickness and wall-height of the nanoengine are 3, 0.3, 0.3 and 1.5\u00a0\u00b5m, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure17-1.png", + "caption": "Fig. 17. Distribution of aspect ratio at the cross section of the SEMMT-applied part.", + "texts": [ + "5 cycles, respectively, which are described by the blue points in Figs. 14 and 15. In Type 3, the deterioration of the mesh began at early stage. The analyses with Type 3 stopped at about 3.5 cycles. On the other hand, in Type 1, the maximum aspect ratio was maintained at about 7. Fig. 16 visualizes the SEMMT-applied part in the analysis of Motion A with Type 1. It can be seen that the SEMMT-applied part deformed along with the deformation of the wing. Therefore, the shape of the SEMMT-applied part is kept as the extension of the shape of the wing. Fig. 17 shows the distribution of aspect ratio. The cross section of the SEMMT-applied part at the middle of the leading edge is shown. At different moments, the similar colored pattern can be seen, which means the mesh pattern in the SEMMT-applied part was maintained. The high rigidity is given in the SEMMT-applied part in Type 1 since the SEMMT stiffening constant C(e) is set to 1000. Therefore, the mesh pattern in the SEMMT-applied part was kept well and the analysis with Type 1 was stable. On the other hand, Type 2 does not have the property to protect the mesh pattern because C(e) is set to 1, which led to the failure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure3-1.png", + "caption": "Figure 3. Crawler track assembly interacting with the ground.", + "texts": [ + " Since the crawler shoe fatigue failure which is the ultimate objective of the present study mainly occurs due to the track-formation interaction and due to the distribution of shovel weight and dipper payload on the crawler shoes [3], the contact forces generated due to interaction between track and other crawler system components can be ignored. However, the shovel load due to the weight of the attachments (dipper, boom and crowd), carbody with swing gear and tapered rollers, propel system, crawler system parts (rollers, driver tumblers, guide rails, idlers, link pin and crawler frame) and weight of extracted material will become part of the distributed load on the tracks which will be included as external forces for crawler dynamic analysis. The multi-body system of the crawler track assembled from crawler shoes is shown in Figure 3. This model consists of 13 crawler shoes and 1 flat flexible rectangular oil sand terrain made up of 50 oil sand units fixed to the default ground link using a mass-spring-damper system from Frimpong and Li [10]. The oil sand terrain has total length, width and depth of 70, 32.5 and 2 m, respectively. The total number of bodies in the system, nc is D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 63. The crawler shoes and oil sand units are also given part numbers to identify them individually as shown in Figure 3. The crawler shoe material is assumed to be steel and their density, volume and mass are listed in Table 2. The density, mass and volume of the oil sand units that makes up the oil sand environment are also listed in Table 2 from Frimpong and Li [10]. The virtual prototype of crawler\u2013terrain model shown in Figure 3 is developed in MSC Adams in Section 3 to simulate the crawler kinematics. In MSC Adams, the location (x, y and z) and orientation of each part in the crawler track assembly are by default identified with respect to the global coordinate system O shown in Figure 3. Body-fixed 313 Euler angles (\u03c8, \u03b8, \u03d5) are used to define the angular orientation of the body relative to the global coordinate system [13]. Also, in MSC Adams, the mass and mass moments of inertia of each body in the system are used both in the kinematics and dynamic analysis of the crawler track assembly [13]. If the system degree of freedom (DOF) is zero, MSC Adams calculates kinematic quantities at each time step and uses dynamic equation of motion with constraint forces to calculate reaction and driving forces due to joints and driving constraints", + " The developed virtual prototype model of crawler kinematics can be extended for further studies to develop dynamic models with flexible crawler shoes to predict shoe failure due to shovel loads and joint link pin loads. All the kinematic equations described below follow the formulation described in Shabana [15,16] for multi-body systems. The crawler shoes in the simplified crawler track model are assumed to be rigid. From Shabana [15], the configuration of each component i in the system is identified using the absolute Cartesian generalised coordinates vector ui. ui \u00bc \u00bdBT i cTi T (1) Bi \u00bc \u00bdBxi ; Byi ; Bzi T and ci \u00bc \u00bdwi; hi; /i T (2) i = 2, 3, 4, \u2026, 64 refers to the body number shown in Figure 3; Bi is the global position vector of the origin of the body-fixed centroidal coordinate system defining the translation of centre of mass of body i; and ci is the set of independent Euler angles that describe the orientation of body i. From Equation (1), the multi-body system shown in Figure 3 has n = 6 \u00d7 63 = 378 generalised coordinates. The generalised coordinates of the system in vector form is given by Shabana [15] as Equations (1) and (2). u \u00bc \u00bdBT 2 cT2 BT 3 cT3 BT 4 cT4 . . . BT 64 cT64 \u00bc \u00bduT2 uT3 uT4 . . . uT64 T (3) The generalised coordinates in Equation (3) are not independent because of the joint and driving constraints [15]. The kinematic constraints equations that describe the joints and motion can be expressed as Equation (4). K u; t\u00f0 \u00de \u00bc K1 u; t\u00f0 \u00de K2 u; t\u00f0 \u00de K3 u; t\u00f0 \u00de Knk u; t\u00f0 \u00de\u00bd T\u00bc 0 (4) K is a vector of constraint functions; nk is the total number of constraint equations for the system; and t is time", + " D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 Ku\u20acu \u00bc E (6) E \u00bc Ku _u\u00f0 \u00deu _u 2Kut _u Ktt (7) _ui \u00bc \u00bd _Bxi _Byi _Bzi _wi _hi _/i T vector of generalised velocity of body i (8) \u20acui \u00bc \u00bd\u20acBxi \u20acByi \u20acBzi \u20acwi \u20achi \u20ac/i T vector of generalised acceleration of body i (9) Ku \u00bc @K @u Jacobian of the constraint equation \u00f04\u00de (10) Kt and Ktt are the partial derivatives of constraint Equation (4) with respect to t; Kut is the partial differential of Equation (10) with respect to t. The algebraic joint and motion constraint equations for the crawler track assembly and flat terrain shown in Figure 3 are described for the following constraints. In the actual crawler track assembly, the crawler shoes are connected to each other using two link pins as shown in Figure 4. The link pins are made to rotate about the axis of the rotation as shown in Figure 4. This rotation action reduces loading in the shoe lugs and increases the life of the link pin [3]. It can be seen from Figure 4 that each crawler shoes can have only relative rotational motion about the joint y-axis due to the pin joint. The relative rotation about the joint z and x axes, and the relative translation motion along the joint x, y and z axes are constrained and negligible", + " This non-linear motion pulling the crawler track may produce maximum fluctuations in the displacement, velocity and acceleration of different crawler shoes in the track. These maximum fluctuations will exert maximum loadings on the crawler shoes which may impact the fatigue life of the crawler shoe. The translation driving constraint is given as Equation (18). Figure 6. Bodies i and j connected by primitive joint. D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 KTD u14; t\u00f0 \u00de \u00bc Bx14 B0;x14 Z vx(t)dt \u00bc 0 (18) B0;x14 \u00bc 17:998m is the initial global x-location of Part 14 (crawler shoe 13) shown in Figure 3 at time t = 0 and vx t\u00f0 \u00de \u00bc 0:012t2 0:0016t3 ; t1 t t2 0:1 ; t[ t2 (19) The maximum velocity of a large mining shovel is reported by Ma and Perkins [12] as 0.25 m/s. This research limits the maximum velocity of the crawler track to 0.1 m/s. In this type, the crawler shoe 13 is given prescribed translation velocity, vx(t), defined in Equation (19) along the global positive x-direction, as well as prescribed positive rotational velocity (x0 z14 ) about the z-axis of the body-fixed coordinate system shown in Figure 7", + " KRD u14; t\u00f0 \u00de \u00bc cos h14 _w14 \u00fe _/14 x0 z14 \u00bc 0 (20) \u03c814 is the rotation angle about original z-axis; \u03b814 \u2013 rotation about the new x-axis; \u03d514\u2013 rotation about the new z-axis; x0 z14 \u2013 angular velocity about the z-axis of body- fixed coordinate system. x0 z14 can be assumed similar to translation motion in Equation (19) as x0 z14 \u00bc 0:12 t2 0:016 t3 ; t1 t t2 1:0 ; t[ t2 deg=s (21) Due to non-holonomic rotational constraint, the crawler track assembly still has five degrees of freedom (two translations and three rotations) to move freely during the turning motion of the crawler track assembly. A section of flexible oil sand terrain model made of 50 oil sand units (bodies 15\u201364 in Figure 3) connected by spherical joints is shown in Figure 8(a) from Frimpong and Li [10] and Frimpong et al. [24]. Each oil sand unit has four motion constraint equations D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 (two translational and two rotational) at its body fixed to centroidal coordinate system that allow only 2 DOFs. This 2 DOFs oil sand unit can be represented as a simple spring-mass-damper system as shown in Figure 8(b) from Frimpong and Li [10] and Frimpong et al", + " KR ui\u00f0 \u00de \u00bc _hi coswi \u00fe _/i sin hi sinwi _wi \u00fe _/i cos hi \u00bc xxi xzi \u00bc 0 (23) KR ui\u00f0 \u00de \u00bc _hi sinwi _/i sin hi coswi _wi \u00fe _/i cos hi \u00bc xyi xzi \u00bc 0 (24) xxi , xyi and xzi are the angular velocities about the global x, y and z axes shown in Figure 9. The oil sand unit 50 (Part 64) is also provided two translation constraint equations given by Equation (25) in addition to two rotational constraints described in Equations (23) or (24). KT u64\u00f0 \u00de \u00bc Bx64 B0;x64 By64 B0;y64 \u00bc 0 (25) D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 B0;x64 \u00bc 66:5m and B0;y64 \u00bc 31:5m are the initial global x and y positions of oil sand unit 50 (Part 64) shown in Figure 3 at time t = 0. In the multi-body system (crawler track + terrain) shown in Figure 3, there are (i) 12 spherical joint constraints due to the connection between crawler shoes (ns = 36 equations); (ii) 12 parallel axis joint primitive constraints (np = 24 equations); (iii) one driving constraint due to translation motion of the crawler track (nd = 1 equation); (iv) one non-holonomic driving constraint due to rotation motion of the crawler track (nnd = 1 equation); (v) 49 spherical joint constraints (nso = 147 equations) and 36 in-plane joint primitive constraints (nio = 36 equations) due to the connection between oil sand units; (vi) 100 non-holonomic rotation motion constraints on oil sand units (nno = 100 equations); and (vii) two translation motion constraints on oil sand unit 50 or Part 64 (nmo = 2 equations)", + " The dynamic equations of motion that relates the inertia forces, external forces, and joint and driving constraint forces are required to completely calculate numerically the accelerations, velocities and position coordinates of each crawler shoe in the track at any time t. The calculated coordinates, velocities and accelerations should also simultaneously satisfy Equations (4\u20136). Since the system cannot be analysed kinematically due to the number of constraint equations being less than the number of coordinates of the system, a virtual prototype model of the crawler track assembly shown in Figure 3 is built in MSC Adams to simulate the crawler\u2013terrain interaction for the sole purpose of capturing crawler kinematics during propelling motion. The virtual prototype of crawler track is built using the following four steps. As a first step, the simplified crawler track assembly shown in Figure 3 is modelled in Solidworks 2013 and the solid model is imported into MSC Adams. The parts of the solid model are then assigned density as listed in Table 2 to define their mass properties. Secondly, the crawler shoes and oil sand units are connected to each other using spherical and primitive joints. The stiffness (k) and damping D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 (c) characteristics of oil sand terrain listed in Table 3 are defined using spring damper element", + " This is because Part 8 which is initially resting on both oil sand units 7 and 12 (Parts 21 and 26) at the beginning of the propelling motion moves completely to oil sand unit 12 (Part 26) at t = 5.3 s as shown in Figure 11(c). It can also be seen from Figure 11(c) that Part 2 in comparison with other crawler shoes exhibits large unsteadiness in z-displacement during the entire propelling motion. This is due to the lack of any joint or motion constraints on one of the sides of crawler shoe 1 (Part 2 in Figure 3). The analytical solution for the velocity of Part 14 in the global x-direction can be obtained by differentiating Equation (26) with respect to time as Equation (27). _Bx14 \u00bc 0:012t2 0:0016t3 ; 0 t 5 s 0:1 ; t[ 5 s m/s (27) Figure 12. Velocity of different crawler shoes in the global x, y and z directions. D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 Equation (27) matches with Adams x-velocity results for Part 14 plotted as a function of time shown in Figure 12(a)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001741_eleco.2015.7394513-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001741_eleco.2015.7394513-Figure3-1.png", + "caption": "Fig. 3. Torques and thrus", + "texts": [], + "surrounding_texts": [ + "All motor on the quadrotor contribute to relative to their angular velocities. It is pos motor produces the force , which is square of the angular speed and the total thr individual thrusts. This is given below constant and is the speed of motor Mi. ! \"# $ %% & $ %% ' ! & The torques that act about the roll and given below. () * + ! (, + - The torque about the yaw axis is diffe above, and expressed as follows where b is t (. / + -! 0 !! + ! 0 (# $()(,(.& 1 *! + ! ! + !/ + -! 0 !! + !" + ] + }, + { + "image_filename": "designv11_34_0001571_2013.15196-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001571_2013.15196-Figure6-1.png", + "caption": "Figure 6. Hohenheim concept of an electrically driven threshing cylinder.", + "texts": [ + " A disadvantage of this solution occurs if additional gearing to adapt the rotational speed of the electric motor and the threshing cylinder is necessary. In this case, the gearbox has to be integrated between the rotor i.e. the casing of the electric motor and the threshing cylinder. The speed adjustment is controlled by an inverter. Through this concept, other rotating devices of combine harvesters such as the rotor of the chopper, the cleaning fan, the beater, the platform auger, or the reel can also be powered. The same can also be applied to a concept originating at the University of Hohenheim for a rotating cylinder (Figure 6) based on the previously mentioned John Deere disclosure. The main difference between the two concepts is that the Hohenheim cylinder uses an electric motor with an inside rotor and an outside stator. The electric motor is completely separated from the cylinder which opens up the possibility for a modularly designed threshing cylinder. The three main modules would be the cylinder, the gear box, and the electric motor. Whereas the design of the cylinder is dependent on the setup of the whole combine harvester, the electric motor can be adapted to the desired capacity of the machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000877_20140824-6-za-1003.01747-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000877_20140824-6-za-1003.01747-Figure10-1.png", + "caption": "Fig. 10 Servomotor with hysteresis brake.", + "texts": [ + " In this illustrated example, a design procedure of the doubleloop control structure is summarized as follows, of which the desired bandwidth is beyond 25 rad/s. i. Set the damping ratio c and the designed bandwidth nc , then one has I PC C . ii. Choose KP and KI such that the internal loop is provided with a satisfactory bandwidth, which is about 5 times the external loop. iii. Plot the dynamic stiffness of MRFF and MBDA; then, choose for setting the bandwidth of . To show the effectiveness of the proposed structure and its design methodology, a rotary servomotor is adopted for the experimental study. As shown in Fig. 10, the servomotor is connected with the hysteresis brake system via a flexible coupling. Therefore, the external disturbance can be realized by controlling the hysteresis brake. A simplified model of the servomotor without the mechanical coupling is obtained with parameters the same as in the illustrated example. Therefore, for the designed bandwidth of velocity loop 30nc rad s and 0.866c , the controller parameters are given by 0.053IC , 0.003PC , 0.0175PK , 1.32IK , and 0.01 . Figure 11 shows the corresponding response of each framework without the mechanical coupling and brake system, in which the input commands are 30rpm and 2000rpm, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001732_j.proeng.2015.12.561-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001732_j.proeng.2015.12.561-Figure4-1.png", + "caption": "Fig. 4. Experimental setup.", + "texts": [ + " Experimental Procedure Synthetic measurements have been generated to check the calibration method behaviour. Although synthetic data tests show a good accuracy improvement, it is not possible to know if this will work under real working conditions. An experiment has been made to check the calibration procedure. A set of 17 reflectors have been placed on a CMM and have been measured by a LT from 5 different positions. Reflector positions have also been measured with the CMM to calculate the initial errors, see Fig. 4. The reflectors were measured with the CMM. These measurements are considered the nominal measurements in the data acquisition step. The LT then measured every reflector in every position, thus obtaining data measurements. The objective function minimizes the differences between all nominal distances given by every pair of reflectors (measured by the CMM), dCMMi, and the same distances measured by every LT, dmik. The number of distances calculated for every position of the LT having 17 reflectors is 136" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000481_ls.1256-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000481_ls.1256-Figure2-1.png", + "caption": "Figure 2. Apparatus photograph: (1) motor, (2) coupling, (3) rotary shaft, (4) inlet and outlet, (5) plexiglass bearing, (6) long-work-distance microscope, (7) CCD, (8) laser, (9) light guide arm, (10) water tank, (11) water pump, and (12) computer.", + "texts": [ + " The difference of this visualisation system compared with a classical PIV was the long-work-distance microscope; its model is QM100 from Questar at the USA. The work distance of this microscope was from 15 to 35.5 cm. The radiuses of field of view were 0.41 mm at work distance 15 cm and 2.29 mm at work distance 35.5 cm. So the field of view was small enough Copyright \u00a9 2014 John Wiley & Sons, Ltd. Lubrication Science (2014) DOI: 10.1002/ls and the work distance was long enough for bearing system. Image-processing software was Dynamic Studio 3.3 from Dantec Dynamics A/S at Denmark. The photograph of this experimental apparatus was presented in Figure 2. The test bearing was presented in Figure 3. The diameter of rotary shaft was 40 mm and the radial clearance was 50 \u03bcm. The external length and width of bearing bush were 80 mm. There were two circumferential recesses on the two sides of bearing for equipping O-shape seal, respectively. The diameter of inlet orifice locating at the bearing centre was 2 mm. The diameter of circumferential recess connecting with inlet was 42 mm, and its axial width was 2 mm. The diameters of these two Copyright \u00a9 2014 John Wiley & Sons, Ltd" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002303_s12221-016-6225-1-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002303_s12221-016-6225-1-Figure6-1.png", + "caption": "Figure 6. Schematic diagram of calibration factors; (a) target levels of calibration and (b) hole ease.", + "texts": [ + " The calibration process was verified over six types of models. Four models were printed for each type with the (Diameter at bottom point \u2212 Diameter at top point) \u00d7100% (1) Diameter at top point diameter of 10, 12.5, 15 and 18 mm. Therefore a total of 72 different models were printed for the verification of calibration process. Models were divided into three groups of circle, cross and the others. Calibration was performed by multiplying certain calibration factors to the target diameters at two levels of 0.5 mm and 8 mm as shown in Figure 6(a). Due to the flow of molten polymer, the bottom of a printed object becomes larger than the initial design. If an object is designed a little bit smaller by multiplying these calibration factors to the intended size, it would be printed in correct dimension. In case of hollow objects, inner diameter was appropriately scaled considering the hole ease as shown in Figure 6(b). The calibration factors were determined based on the relationship between flow and rheological properties. The temperatures of extrusion nozzle and heating bed were determined based on the DSC analysis results. It was found that the diameter at top point was proportional to the weight percentage at extrusion temperature. The value of hole ease could also be determined based on this result. The swelling ratio was determined by the result of viscometry. Prediction of Swelling Ratio Calibration factor at each level was the inverse of the swelling ratio at that level" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002025_s12541-015-0245-4-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002025_s12541-015-0245-4-Figure1-1.png", + "caption": "Fig. 1 Determination of the configuration of a two-identical-link robot", + "texts": [ + " But the L2-GVG is not necessarily connected, and we have used the L2R-edge to connect dispersive L2-GVG edges to produce the roadmap L2-HGVG, where the robot is tangent to a planar GVG structure with the same orientation of the two rods. The L2 robot is composed of two rods of equal length L with a revolute joint P between them. The points Q1 and Q2 are the endpoints of the rod 1 and rod 2. The configuration q = (p, \u03b81, \u03b82) \u2208 R2 \u00d7 T2 of the L2 robot is described by the position vector p of the joint point P and the orientations \u03b81 and \u03b82 of the two rods with respect to the horizontal in Fig. 1. Let Ra(q) denote the set of points in R2 occupied by the a-th rod of the L2 robot at configuration q. Then Ra(q) can be written as (1) The L2-HGVG is defined using equidistance relationships based on a distance function. The distance is the workspace distance between robot Ra and obstacle Ci: (2) where the subscript i (i \u2208 {1, 2, \u2026, n}) is used to number the index to the obstacle. Note the Ci is the convex obstacle, concave obstacles can be modeled as unions of convex obstacles. The distance gradient is represented as (3) where ra and ci are the nearest points on the rod a and obstacle Ci, respectively, and \u03b4ij is the Kronecker delta: (4) If the rod Ra is parallel to the surface of an obstacle, the nearest points are continuous and infinitely many" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001732_j.proeng.2015.12.561-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001732_j.proeng.2015.12.561-Figure1-1.png", + "caption": "Fig. 1. Laser tracker kinematic model", + "texts": [ + " As these data have been generated according to our model, the error parameters calculated give a very good LT accuracy, so it is necessary to use real data to know the accuracy increment our calibration method can obtain. The calibration procedure consists of identifying the geometric parameters in order to improve the measurement model accuracy. This process can be carried out in four steps: determination of the kinematic model by means of non-linear equations, data acquisition, geometric parameter identification and model evaluation. 2. Methodology A kinematic model of the LT based on the Denavit-Hartenberg method [6] has been developed, see Fig. 1. The kinematic model establishes mathematical relations and obtains non-linear equations that relate the joint variables with the position and orientation of the end-effector [7]. This method has been widely used in mechanism modelling [8-10], and uses four parameters to model the coordinate transformation between successive reference systems. This method models the coordinate transformation between successive reference systems, using four parameters (distances di, ai, and angles \u03b8i, \u03b1i). The homogenous transformation matrix between frame i and i-1 depends on these four parameters as shown in equation 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003748_iecon43393.2020.9254887-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003748_iecon43393.2020.9254887-Figure8-1.png", + "caption": "Fig. 8. Configuration of studied outer rotor IPMSM. (a) Type III (b) part structure of Type III", + "texts": [ + " DESIGN OF ROTOR SLOT In this section, a new rotor slotting design method was presented to reduce torque ripple for outer rotor IPMSM. According to the basic concept in section II of the article, we can know that rotor slots can modify the gap flux density, and reduce the cogging torque. Type III model has been analyzed by FEA, which takes the outer-rotor permanent magnet motor for example that is reliable and suitable for high speed operation. Combined with theoretical analysis, a new rotor slotting design method was presented to reduce torque ripple. Fig.8 shows the configuration of the type III outer rotor IPMSM. To be more comparative and persuasive, the structure of type III is the same as the structure of type II except for the rotor slots. The type III model in Fig.8 is analysed by FEM. To make the comparison more convincing, all the settings (including the lead angle, motor speed, input voltage, winding settings, the shape of the flux-barrier and etc.) are the same except the rotor slots. The cogging torque of type II and type III are shown and compared in Fig.9. From this FEM analysis it can be noticed that the peak value of the calculated cogging torque is equal to 0.06 Nm for the type II model, and it is equal to 0.045 Nm for the type III model. Through the simple numerical calculation, we can get that, the cogging torque of type III model is 25% lower than that of type II model. The graph in Fig.10 is the torque of the type III model shown in Fig.8. In this part, output torque, torque ripple and induced voltages of type II and type III are analyzed and compared. 2692 Authorized licensed use limited to: Carleton University. Downloaded on May 31,2021 at 22:30:50 UTC from IEEE Xplore. Restrictions apply. From Fig.10, we can find that the average torque of type III is 1.75 Nm, which is the same as the type\u2161. Through the simple numerical calculation, we can get the torque ripple of the Type III is 11.43%, while the torque ripple of the Type\u2161 is only 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001053_ecc.2014.6862509-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001053_ecc.2014.6862509-Figure7-1.png", + "caption": "Fig. 7. The well known cart and pendulum system.", + "texts": [ + " The reason is the feature rich 32bit Flash CortexM4 based MCU very well fitted for general control purposes: \u2022 up to 180 MHz/225 DMIPS, with DSP instructions, floating point unit (FPU; single precision) and advanced peripherals \u2022 2 DAC 12bit, 3x 12bit ADC (24 channels) \u2022 synchronized PWM-timers, quad-encoder channels \u2022 available discovery-board (approx. 15$; programmer and debugging probe already on-board; just requires USB and serial (for scope) connection) Towards industrial targets (Programmable Logic Controller, PLC) there is ongoing development concerning the application to a b&r Powerpanel (PP400) using the IDE approach with the target specific Automation Studio. As a reference application the well known cart and pendulum system in gantry crane configuration is presented, see Fig. 7. As common to real world crane applications at least the equivalent length to the center of mass of the load is unknown. Instead, we assume the equivalent length of the pendulum rod l2 as unknown but constant. The cart with mass m1 is driven by a dc drive with external excitation, see Fig. 8. In the following, we assume a very small electrical time constant (\u03c4el = LA RA , with armature inductance LA and resistance RA) compared to the mechanical one. By means of system reduction, i.e, LA \u2192 0, we introduce equivalent parameters integrating the whole drive-train (drive constant km, inertia JA, transmission ratio n, gear pinion radius r, and some mechanical damping d1) into the mathematical model of the cart" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000004_iros40897.2019.8968055-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000004_iros40897.2019.8968055-Figure4-1.png", + "caption": "Fig. 4. Walking disturbance experiments.", + "texts": [ + "44% of total virtual work (W\u0304+ v +W\u0304\u2212v ) in FW experiment, and this means virtual spring is almost doing positive work to the hip joint of swing leg throughout the swing phase of walking. This result indicates that the virtual torque \u03c4v generated by the virtual stiffness model can assist the motion of the human hip joint during forward walking. To investigate whether virtual torque is able to assist human to regain balance when they are suffering from a disturbance force, the walking disturbance experiments shown in Fig.4 are carried out next. In the WFP experiment, as shown in Fig.4(a), each subject was asked to walk forward for ten times, and they will suffer from an unexpected disturbance force acting on their chest. As shown in Fig.2(b1), a forward force is acting on the chest of the subject at about 20% of swing phase, and then CP is 8188 Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 04,2020 at 14:12:08 UTC from IEEE Xplore. Restrictions apply. accelerated and exceeds the upper boundary of BoS (ToeX) at 35.9% of swing phase. Human finally regains balance at 86.6% of the swing phase. Fig.2(b2) shows that during 35.9% to 64.4% of swing phase there is \u03c4v \u00b7 THip sw > 0, and as shown in Fig.2(b3), in this time intervals Pv is an \u2018assisted power\u2019. Moreover, Table I shows that W+ v accounts for 92.81% of total virtual work in WFP experiment, and this result shows virtual torque can assist hip joint to regain balance when the human suffers from a forward disturbance force during walking. In the WBP experiment, as shown in Fig.4(b), subjects will suffer from an unexpected disturbance force acting on their back when they are walking forward. Each subject repeats the experiment ten times. As we can see from Fig.2(c1), human suffers from a backward disturbance force at about 41% of swing phase and then CP begins to move backward. At the 63.3% of swing phase, CP exceeds the lower boundary of BoS (HeelX) and is caught up by swing foot at the 96.9% of swing phase. Fig.2(c2) shows \u03c4v and THip sw are in the same direction during 63" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003248_s00170-020-05597-z-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003248_s00170-020-05597-z-Figure7-1.png", + "caption": "Fig. 7 Verification experiment of ball screw joint", + "texts": [ + " The parameters of the guide are listed in Table 2. The material of the fixed joint is no. 45 steel, and the contact surfaces are machined by grinding. The contact surface parameters were measured by atomic force microscope of Bruker Company and identified by the structure function method [21, 24]. The preload of the surfaces is controlled by the tightening torque of the bolts. The parameters of the fixed joint are listed in Table 3. The verification experiment of the ball screw joint is shown in Fig. 7. As shown in Fig. 7a, the ball screw was suspended by flexible ropes, and the nut was moved to the middle position of the screw. The frequency response functions obtained by different excitation positions in the modal test are shown in Fig. 7b. According to the calculation parameters in Table 1, the predicted axial stiffness of the ball screw joint is 9.02 \u00d7 108 N/m. The finite element model is shown in Fig. 7c, and the axial stiffness of the ball screw joint is entered into the BUSHING element. The comparison between theoretical and experimental values for the ball screw joint is shown in Table 4. In addition, some experimental stiffness values of ball screw joints under certain conditions have been given by the THK Company: First, apply the 0.1Ca preload to the ball screw joint and then apply 3 times the preload as the axial force to the ball screw joint. Finally, the stiffness value is obtained according to the axial displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003888_s11771-020-4551-3-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003888_s11771-020-4551-3-Figure2-1.png", + "caption": "Figure 2 Dynamic model of gear system", + "texts": [ + " rbp and rbg denote the radii of the base circles of the gears (p represents the pinion gear and g denotes the driven gear), and their angular velocities are denoted as \u03c9p and \u03c9g, respectively. For high-speed and heavy-duty gear systems, J. Cent. South Univ. (2020) 27: 3350\u22123363 3352 the inertial force and damping have significant effects on the transmission characteristics, and thus, dynamic analysis is necessary. Considering the effect of friction, a dynamic model of the gear system is established, as shown in Figure 2. The model takes the torsional and translational vibrations of spur gears into consideration. As shown in Figure 2, the y-axis is in the direction of the LOA ( ).AD The torsional angular displacements of the pinion and driven gears are denoted as \u03b8p and \u03b8g, respectively. Ip represents the moment of inertia of the pinion gear, and Ig represents the moment of inertia of the driven gear. The meshing stiffness and damping are denoted as kt and ct, respectively. The static transmission error is et. For a standard gear transmission, at the meshing point K, the linear velocity of the pinion and driven gears should be equal", + " The friction is in the positive direction of the x axis if \u03bbi>0, where 1 p b b b 2 b c sgn( ), 0 sgn[mod( , ) ], 0 0, ct t t t t t t t t t t t = = (4) Based on the dynamic model of the gear system, the differential equation is established as [39]: p p p bp p g g g bg g p p p p p p p g g g g g g g p p p p p p fp g g g g g g fg ( ) ( ) ( ) ( ) ( ) ( ) ( ) 0 ( ) ( ) ( ) 0 ( ) ( ) ( ) 0 ( ) ( ) ( ) 0 - - y y y y x x x x I t T r F t I t T r F t m y t c y t k y t F m y t c y t k y t F m x t c x t k x t F m x t c x t k x t F + + = + = + + = + = (5) where xp(t) and xg(t) represent the displacement of the pinion gear and driven gear caused by vibration along the x direction, respectively. The Weber method [40] was used to determine the time-varying meshing stiffness of the gear system. The fourth-order Runge-Kutta method was applied to solve the differential equations, and dynamic characteristics, such as the dynamic load and fluctuating velocity during the gear transmission process, were calculated. 2.3.1 Basic contact parameters of gear system As shown in Figure 2, according to the properties of involute curves, the radii of curvature at any possible meshing point K can be denoted as follows: p 1 1 bp g 2 2 bg ( ) tan ( ) ( ) tan ( ) R K N K N P PK r s t R K N K N P PK r s t (6) where s denotes the distance from the meshing point K to the pitch point P; \u03c6 represents the J. Cent. South Univ. (2020) 27: 3350\u22123363 3353 working pressure angle. The value of \u03c6 can be obtained by the following equation: 1 2 1 2 (inv inv ) 2 tan z z x x x (7) where z1 and z2 denote the tooth number of the pinion gear and driven gear, respectively; and the corresponding modification coefficients are denoted as x1 and x2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003748_iecon43393.2020.9254887-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003748_iecon43393.2020.9254887-Figure1-1.png", + "caption": "Fig. 1. structure of the outer rotor IPMSM", + "texts": [ + " Section III describes the technique to reduce the torque ripple by promoting the design of flux-barrier, and verify it by FiniteElement Analysis (FEA). The technique based on rotor slotting design is introduced in section IV. The operation performance of the motor before and after optimization is compared and analyzed in section V. Finally, in section VI, concluding remarks are made. II. BASIC CONCEPT AND METHOD OF ANALYSIS An outer rotor IPMSM consists of rotor, stator, permanent magnets, windings, flux-barrier and shaft. Fig.1 shows the shape and structure of the outer rotor IPMSM, which has 8 poles and 24 slots with distributed windings. 978-1-7281-5414-5/20/$31.00 \u00a92020 IEEE 2689 Authorized licensed use limited to: Carleton University. Downloaded on May 31,2021 at 22:30:50 UTC from IEEE Xplore. Restrictions apply. D-q reference frame is always adopted to calculate the current and voltage of outer rotor IPMSMs for computing simplifying. The direction of the rotor magnetic field is taken as the d-axis, which is on the center line of the rotor\u2019s magnetic pole" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000584_j.jsc.2014.09.031-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000584_j.jsc.2014.09.031-Figure1-1.png", + "caption": "Fig. 1. A regular pentagon obtained by the simplest knot of three folds.", + "texts": [ + " We employ a software tool called Eos (e-origami system), which incorporates the extension of Huzita\u2019s basic fold operations for construction, and Gr\u00f6bner basis computation for proving. Our study yields more mathematical rigor and in-depth results about the polygonal knots. \u00a9 2014 Elsevier Ltd. All rights reserved. Given a rectangular origami, i.e. sheet of folding paper of adequate length, we can construct the simplest knot by performing three times of folds. The shape of the knot is made to be a regular pentagon if we fasten the knot tightly without distorting the origami as shown in Fig. 1. The method is so simple that it should be known since the age of early human civilization. However, only from \u2729 This work was supported by JSPS KAKENHI Grant No. 25330007. The second author of this paper is supported by the postdoctoral fellowship at Kwansei Gakuin University. This is the revised and enhanced version of the earlier papers published in the conference proceedings: Ghourabi et al. (2013a), Ida et al. (2014). E-mail addresses: ida@cs.tsukuba.ac.jp (T. Ida), ghourabi@kwansei.ac.jp (F", + " Further discussions on the adequacy of operations (O1)\u2013(O6) as the set of basic fold operations are shown in Ghourabi et al. (2013b), for example. The above statements of Huzita\u2019s basic fold operations are about the fold lines, and do not address to the issues of the directions of the fold, i.e. mountain or valley fold, or of the choices of the faces to be moved. When we perform a fold in Eos, we need to specify them unless we rely on their default values of Eos, i.e. valley fold and the choice of all the faces to the left of the fold line (interpreted as a vector). We depicted a knotting operation in Fig. 1. In the figure we made one simple knot. When the height of the tape is infinitesimal and both ends of the tape are connected, the tape can be abstracted as a closed curve, i.e. an object of study in the theory of knots. The knot with 3 crossings is the most basic. It is denoted 31 in the Alexander\u2013Briggs notation (Alexander and Briggs, 1926). When the height is finite, each crossing exhibits a certain polygonal shape. The collection of the crossings, some of which overlap each other, is called a polygonal knot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001976_17480930.2015.1093761-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001976_17480930.2015.1093761-Figure9-1.png", + "caption": "Figure 9. Oil sand units with 2 DOFs connected by spherical and in-plane joints [10,24].", + "texts": [ + " Each oil sand unit has four motion constraint equations D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 (two translational and two rotational) at its body fixed to centroidal coordinate system that allow only 2 DOFs. This 2 DOFs oil sand unit can be represented as a simple spring-mass-damper system as shown in Figure 8(b) from Frimpong and Li [10] and Frimpong et al. [24]. However, when oil sand units with 2 DOFs are connected by spherical joints, redundant constraints equations are introduced in the oil sand model. To eliminate the redundancies, the combination of spherical and primitive in-plane joints is used to connect the oil sand units (Figure 9). In addition, two translation motion constraints on the oil sand units 1\u201349 (Parts 15\u201363) are also removed. The constraint equations for spherical joints connecting two adjacent oil sand units shown in Figure 9 are similar in form to Equations (13) and (14) and hence not shown here. The constraint equation for in-plane joint can be expressed as in Equation (22) [13]. KI ui; uj \u00bc rzi rzj \u00bc 0 (22) rzi and rzj are the global z-coordinate of the in-plane joint location point on two adjacent oil sand units shown in Figure 9. The two non-holonomic rotation motion constraint equations on oil sand units (bodies i = 15 \u2013 64) follows either Equation (23) or (24) depending on the orientation of spring damper system shown in Figure 9. KR ui\u00f0 \u00de \u00bc _hi coswi \u00fe _/i sin hi sinwi _wi \u00fe _/i cos hi \u00bc xxi xzi \u00bc 0 (23) KR ui\u00f0 \u00de \u00bc _hi sinwi _/i sin hi coswi _wi \u00fe _/i cos hi \u00bc xyi xzi \u00bc 0 (24) xxi , xyi and xzi are the angular velocities about the global x, y and z axes shown in Figure 9. The oil sand unit 50 (Part 64) is also provided two translation constraint equations given by Equation (25) in addition to two rotational constraints described in Equations (23) or (24). KT u64\u00f0 \u00de \u00bc Bx64 B0;x64 By64 B0;y64 \u00bc 0 (25) D ow nl oa de d by [ U ni ve rs ite L av al ] at 0 2: 01 0 2 D ec em be r 20 15 B0;x64 \u00bc 66:5m and B0;y64 \u00bc 31:5m are the initial global x and y positions of oil sand unit 50 (Part 64) shown in Figure 3 at time t = 0. In the multi-body system (crawler track + terrain) shown in Figure 3, there are (i) 12 spherical joint constraints due to the connection between crawler shoes (ns = 36 equations); (ii) 12 parallel axis joint primitive constraints (np = 24 equations); (iii) one driving constraint due to translation motion of the crawler track (nd = 1 equation); (iv) one non-holonomic driving constraint due to rotation motion of the crawler track (nnd = 1 equation); (v) 49 spherical joint constraints (nso = 147 equations) and 36 in-plane joint primitive constraints (nio = 36 equations) due to the connection between oil sand units; (vi) 100 non-holonomic rotation motion constraints on oil sand units (nno = 100 equations); and (vii) two translation motion constraints on oil sand unit 50 or Part 64 (nmo = 2 equations)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001571_2013.15196-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001571_2013.15196-Figure4-1.png", + "caption": "Figure 4. Hohenheim combine with serial hybrid electric ground drive.", + "texts": [ + " THE ELECTRIC COMBINE HARVESTER To build up a totally or mostly electrically driven combine harvester, it becomes necessary to initially replace the existing mechanical and hydrostatic drives with electrical drives. The second step would be to reconsider the setup of the whole machine to appropriately use the advantages of the electric system. Ground drive Concerning the first step, some approaches have already been made. For example, a combine used at the University of Hohenheim is equipped with a serial hybrid electric ground drive using 650 VDC for the power circuit and has a standard voltage of 400 V for the three-phase AC machines (Figure 4). For use as generators in the hybrid wheel drive system, permanent magnet synchronous machines are preferable to asynchronous machines due to their higher power to weight ratio (Barucki, 2001). Therefore, these machines are already used for most serial hybrid applications. For the wheel drive itself, permanent magnet synchronous machines with outside rotors seem to be most promising since they have a low weight and a high starting torque compared to other electric machines (Weck et al., 2000). If a machine is designed in Germany, the DIN standards set by the German Institute for Standardization have to be observed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure25-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure25-1.png", + "caption": "Fig. 25. Geometry of the torus and coordinates \u03be , \u03b7.", + "texts": [ + " Nevertheless, enhancement of the mathematical model can be carried out considering the wheels as two tori instead of two rings. This is particularly important in the case of the Waveboard, where the major and minus radius of the torus are more similar than in other systems like a bicycle. In order to define the toroidal surface, two angular coordinates are required, as opposed to the ring case, where only the non-generalized coordinate \u03be i was necessary. Therefore, the angular coordinates \u03be and \u03b7 for toroidal and poloidal directions are defined. Fig. 25 shows the geometry of the torus, with b being the major radius (distance from the center G of the torus to the center of the tube P ), and a the radius of the tube. The first step to construct the model is locating the centers of the bodies 4 and 7 (denoted by G 4 and G 7 ), whose absolute position vectors can be found in Appendix A. Starting from body frames 4 and 7 (for the rear and forward wheel, respectively), one can define a counterclockwise rotation \u03be i about the Y 4 , Y 7 axes in order to orientate axes X \u03beR and X \u03beF , so that they point to the center of the tube P " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001822_s40430-015-0458-6-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001822_s40430-015-0458-6-Figure1-1.png", + "caption": "Fig. 1 The five zones of interest, marking the five different fault types", + "texts": [ + " This problem falls within the area of fault detection and diagnosis that aims to automate the process of discovering faults and diagnosing their causes in mechanical engineering systems by extracting the relevant information from raw data [2]. There are many causes for rolling bearing faults, such as improper mounting, acid corrosion, bad lubrication, wrong design and plastic deformation [3]. However, the most common defect is cavities or cracks on the surfaces of the bearing elements, caused by periodic loads. This is known as material fatigue, and it affects mostly the inner and the outer bearing race (in 90 % of the cases), while the rest 10 % affects the rolling elements (balls in our case) and the cage [4] (Fig. 1). A common way to detect such faults is by analyzing vibration signals during operation, as recorded by accelerometers. Considerable analysis has been conducted regarding this issue, as described by Abbasion et al. [1], Purushotham et al. [2], Wang and McFadden [5], Shiroishi et al. [6], Scholkopf [7], Dellomo [8], Li et al. [9], Jack and Nandi [10], Nikolaou and Antoniadis [11], Samanta et al. [12], Al-Ghamd and Mba [13], Trendafilova [14], de Moura et al. [15], etc. A number of computational intelligence (CI) methods are reported to be used in the above-mentioned literature, namely support vector machines, artificial neural networks, wavelet transform, principal component analysis, genetic algorithms, etc", + " Figure 3 summarizes the data being made available. Outer raceway faults are stationary faults, therefore placement of the fault relative to the load zone of the bearing has a direct impact on the vibration response of the motor/bearing system. In order to quantify this effect, experiments were conducted for both fan and drive end bearings with outer raceway faults located at 3 o\u2019clock (orthogonal to the load zone), at 6 o\u2019clock (directly in the load zone), and at 12 o\u2019clock (opposite to the load zone (Fig. 1) [22].Fig. 2 Representation of the test stand Fig. 3 Faulty bearing data 1 3 Let A = {(xi, yli), i = 1, . . . ,N , l = 1, . . . , L} be a given set of data instances consisting of input vectors (observations) xi, and a target value for classification, yl, for a number of N timesteps and L possible classes. The use of an aggregation function Af allows for mapping the data instances (input space) to the feature space F, i.e. where a is the aggregation measure and i = 1, \u2026, N. The aggregation measure a represents a population of xn instances, = 1, \u2026, a, from which a number of k = 1, \u2026, m statistical parameters (features) is calculated for the time domain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001240_aim.2014.6878315-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001240_aim.2014.6878315-Figure3-1.png", + "caption": "Fig. 3. A planetary gear set with four planets", + "texts": [ + " The planetary gearbox used in this work has two identical planetary gear sets in a back-to-back configuration such that the overall speed ratio is 1/4\u00d74 =1.0, where 4 is the speed ratio for each gear set. The sun gear of one gear set is connected to the driving motor, while the sun gear of the other gear set is connected to the load generator. The carriers of the planetary gear set are connected together through a steel spline, and the annulus gears are fixed to the housing. Each planetary gear set consists of four equally spaced planets as illustrated in Fig. 3. A multi-DOF lumped parameter model for a planetary gear set is developed and can be expressed in matrix form [14] as jjjjgjj \u03a4QKQCQM =++ &&& (4) where j = 1 or 2 represents one of the two gear sets, Mj is the mass matrix, Cj is the damping matrix, Kj is the stiffness matrix, and Tj is the external applied torques vector. Since only the torsional DOF is significant in modeling the gearbox coupling with the motor, the vector Qj is chosen to contain all the rotational DOFs of the gear set as [ ]Txjjpjpjpjpsjj \u03b8\u03b8\u03b8\u03b8\u03b8\u03b8 4321=Q (5) where s represents the sun gear, x represents the carrier and p1 to p4 represent planet gears 1 to 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002485_j.triboint.2016.10.041-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002485_j.triboint.2016.10.041-Figure5-1.png", + "caption": "Fig. 5. Thermocouple positions.", + "texts": [ + " The collar has been covered with an aluminum case, with the aim of covering but also containing the circuital modules, even during the rotation, to prevent collapses due to the centrifuge force. The assembly of the collar and its case has been keyed on the shaft (Fig. 4). Once the circuit was placed into the collar and covered by the aluminum case, the thermocouple probe has been placed in contact with the inner portion of the rotating ring through a hole, drilled in the rotating ring housing and in the rotating counter-face of the seal, as shown in Fig. 5. The thermocouple probe has been introduced into the hole and, taking into account the high centrifuge forces during the rotation, a reliable contact between the probe and the ring has been assured. Another thermocouple is connected to the stationary ring. The validation of the new wireless thermometric chain has been performed by two different procedures: one using a hot plate and another using the tribological test rig, similarly to Gupta et al. in [8]. In both cases a reference thermometric chain has been assembled in parallel with the wireless chain and connected to a same specimen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001376_ecce.2013.6647067-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001376_ecce.2013.6647067-Figure6-1.png", + "caption": "Fig. 6. Experimental Environment", + "texts": [ + " The dq/3-phase transform matrix Cdq UV W at \u03b8e = 0 is given by Cdq UV W = \u221a 2 3 \u23a1 \u23a2\u23a3 1 0 \u2212 1 2 \u221a 3 2 \u2212 1 2 \u2212 \u221a 3 2 \u23a4 \u23a5\u23a6 . (22) This means the flux by q-axis current pass only through V and W-phase teeth. In Fig. 4, magnetic flux and q-axis flux affect each other, and this leads to asymmetry of flux distribution on V and W-phase teeth. It is difficult to measure radial force directly. In this paper, radial acceleration outside the stator is evaluated in stead of radial force. Experimental environment is shown in Fig. 6. The velocity is controlled by load motor. Drive motor controls current and radial acceleration on the stator of the drive motor is measured by accelerator. In this paper, the negative d-axis current reference is limited within -20 A because of the motor rating. Experimental results of run-up spectrograms with and without id current are shown in Fig. 7. The number of pole pairs are 6 in this motor, so the frequency of 2nd order acceleration is n 5 [Hz], in which n is rotation speed. Fig. 7 shows 2nd and 6th acceleration is relatively larger compared with other components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002799_6.2020-1117-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002799_6.2020-1117-Figure3-1.png", + "caption": "Figure 3 - Cable Kinematics", + "texts": [ + " Figure 2 - System Model (4 Quadcopters) For the initial formulation of this problem, the payload and UAVs are considered as point masses, and so their orientation (and the frames attached to them) is tied to the global frame. Additionally, the UAVs will be initially be considered as omnidirectional robots capable of accelerating in any direction. This model allows us to build off our previous work with GFM directly, before moving on to more sophisticated models. For quadcopters, dynamics and kinematics can be adapted from Vijaykumar et al. [14], and the payload attachment points can be moved out in order to consider rigid body dynamics of the payload. IV. 2D-Kinematic System From Figure 3, the length and orientation of one cable can be found using the kinematic equation: \ud835\udc59\ud835\udc56\ud835\udc57\ud835\udc56 = ?\u20d1?\ud835\udc56 \u2212 \ud835\udc5d (1) If the attachment points are off the center of mass (true in the rigid body case), an extra vector would be added to account for the offset. D ow nl oa de d by A U B U R N U N IV E R SI T Y o n Fe br ua ry 2 , 2 02 0 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 02 0- 11 17 Additionally, from the geometry of the system it can be shown that the rate of change in cable length is represented by the projection of the relative velocity of the connected UAV with respect to the payload onto the cable pointing vector: \ud835\udc59" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure8.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure8.1-1.png", + "caption": "Figure 8.1 A spherical wrist and a nonspherical wrist.", + "texts": [ + " Such simplified forward kinematic equations will be obtained for several serial manipulators in Chapter 9. A manipulator with a spherical wrist is such that its hand is connected to its arm with three successive revolute joints, whose axes intersect each other at a single point R, which is defined as the wrist point. Here, as usual, the hand is defined as the part of the manipulator beyond the wrist point and the arm is defined as the part of the manipulator up to the wrist point. The end-effector is integrated with the last link. A spherical wrist and a nonspherical wrist are illustrated in Figure 8.1. The wrist on the left-hand side is spherical because the axes of the joints m, m\u22121, and m\u22122 intersect each other at the same point R, which is the wrist point as defined before. On the other hand, the wrist on the right-hand side is nonspherical because the axes of the jointsm\u22121 and m\u22122 do not even intersect each other. Yet, for a nonspherical wrist, the point R, which happens to be the center of the penultimate revolute joint m\u22121 that precedes the tip point P, is still defined as the wrist point", + "73) constitute a complete set of m* = m+ 6 scalar equations that contain m* = m+ 6 unknowns, which are the m joint variables (q1, q2, \u2026, qm) and the six Lagrange multipliers (\ud835\udf061, \ud835\udf062, \ud835\udf063; \ud835\udf06\u20321, \ud835\udf06 \u2032 2, \ud835\udf06 \u2032 3). Thus, these m* scalar equations can be solved to find the optimal values of the joint variables together with the corresponding values of the Lagrange multipliers. A serial manipulator is defined to be deficient if m < \ud835\udf07 \u2264 6 (8.74) For a deficient manipulator with a regular end-effector such that u\u20d7(m) 3 = u\u20d7a as shown in Figure 8.1, the following forward kinematic equations can be written for the tip point location and the end-effector orientation. p = \ud835\udf19(q1, q2, q3,\u2026 , qm\u22123, qm\u22122, qm\u22121) (8.75) C\u0302 = F\u0302(q1, q2, q3,\u2026 , qm\u22123, qm\u22122, qm\u22121, qm) (8.76) With a deficient manipulator, even if the intention is to have \ud835\udf07 for the task mobility, it is not possible to realize this intention. Therefore, it is necessary to have a compromise by reducing the required task mobility from \ud835\udf07 to \ud835\udf07 \u2032 = m. This compromise makes the manipulator regular and thus the joint variables can be found so as to satisfy the \ud835\udf07\u2032 specifications (instead of \ud835\udf07) for the position (location and orientation) of the end-effector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003174_j.matchar.2020.110401-Figure29-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003174_j.matchar.2020.110401-Figure29-1.png", + "caption": "Fig. 29. (a) Schematic diagram of the wear mechanism of microscratch in HT1 and HT2 conditions and (b, c, d) schematic illustration of the microscratch at the interface matrix/carbides hard phase.", + "texts": [ + " 28 shows that under the high load, the main defects in the scratches are deep grooves, burrs formed by extrusion accumulation and microcracks. The scratch widths of the UT alloy and those that underwent HT1, HT2, HT3 under a load of 1000 mN are 19 \u03bcm, 28 \u03bcm, 27 \u03bcm and 19 \u03bcm, respectively. The results reveal that the NiCrSiB +0.5 wt% Y2O3 alloy samples prepared by LAM present excellent wear resistance after HT3. Due to the discrete distribution of the Cr-carbide hard phases, the microwear properties of the samples subjected to HT1 and HT2 were poor. Therefore, the microscopic wear mechanism is illustrated by a schematic diagram. Fig. 29 shows a schematic diagram of the wear mechanism of the microscratches in the HT1 and HT2 samples. As shown in Fig. 29a, the Cr-rich hard carbides, marked in yellow, were discretely distributed on the surface of the test sample. When the microindenter was moved from left to right, it came into contact with the Cr-rich hard phase, resulting in a high contact stress. As a result, the scratch width of the hard phase region is narrow, and the fluctuation of the friction coefficient and the scratch residual depth are substantial. Therefore, the samples that underwent HT1 and HT2, in which the hard phase is discretely distributed, have poor wear resistance. Fig. 29(b\u2013d) illustrates the scratching process of the spherical indenter in the Cr- carbide area [28]. As the indenter approaches the hard carbide, it increases the contact pressure at the interface between the matrix and carbide, resulting in the matrix deforming. As the indenter continues to move, the contact stress continues to increase, causing the hard blocky phase to be pressed into the soft matrix, thereby facilitating the indenter through the region with the hard phase. Therefore, this results in a shallow and narrow scratch in the region with the hard phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002834_tdmr.2020.2966043-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002834_tdmr.2020.2966043-Figure1-1.png", + "caption": "Fig. 1. The polyamide scanner, a) manufactured device b) CAD drawing, c) 1st (fundamental) mode shape (out-of-plane) of the device. The cyclic tests were performed via actuating the device in the fundamental mode.", + "texts": [ + " Polyamide is an extremely versatile material in biomedical research (adapted for use in medical devices for encapsulation and insulation of active medical implants, moisture absorption, prevention of corrosion, also exhibiting biostability and biocompatibility) [12]. The paper starts with the structural description of the laser scanners under test and the experimental setup in Section II. Then, results will be provided and discussed in Section III. Finally, conclusions will be drawn in Section IV. The scanner illustrated in Fig. 1, having 10 x 10 mm2 die area, 4-mm long flexure, and 3 x 7 mm2 head area is designed in the shape of hammerhead to allow for simultaneous excitation of out-of-plane and torsional modes to scan a Lissajous pattern. The devices were manufactured using selective laser sintering (SLS) of PA polyamide (using the EOS Formiga P110 instrument). The device is magnetically actuated, ideally through driving an external coil (located behind the scanner) with the sum of both mechanical mode frequencies to excite both modes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000816_s10778-014-0651-9-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000816_s10778-014-0651-9-Figure1-1.png", + "caption": "Fig. 1", + "texts": [ + " Here we will analytically derive, for the first time, conditions for the origin, existence, cessation, and stability of various steady-state motions of an IS consisting of a rotating body and two identical simple pendulums fit on its longitudinal axis and estimate the residual nutation angle. Since pendulum, ball, and fluid (ring) dampers suppress nutation in a similar way [5], such an IS models how these dampers decrease or increase the nutation angle depending on whether they are correctly or incorrectly installed on the satellite. Related issues are addressed in [16\u201319]. 1. Description of the Isolated System. Its Axial Moment of Inertia. The IS consists of a body and two pendulums forming a nutation damper (Fig. 1). The body has center of mass at the point O, mass M, and axial moments of inertia A, B, C about the principal central axes of inertia O . Two simple pendulums of length l and mass m/2 each are fit on the longitudinal axis of the body (Fig. 1b). The pendulums move in the plane \u00ce 1 1 1 parallel to the plane O located at a distance b from it. When the pendulums move relative to the body, they are subject to viscous drag whose magnitude and nature are shown [4] to be insignificant for the problem under consideration. The steady-state motion of the IS is rotation about the z G -axis on which its angular momentum vector and the center of mass G lie. During the primary motion, the pendulums lie on the straight line 1 and the IS rotates about the longitudinal axis of the body, = z G (Fig. 1a). Since the system is symmetric, during the secondary motions, the pendulums make an angle with the straight line 1 (Fig. 1b) and the -axis makes an angle with the z G -axis (Fig. 1c). International Applied Mechanics, Vol. 50, No. 4, July, 2014 1063-7095/14/5004-0459 \u00a92014 Springer Science+Business Media New York 459 Kirovograd National Technical University, 8 Universitetskii Av., Kirovograd, Ukraine 25006, e-mail: fgb@online.ua. Translated from Prikladnaya Mekhanika, Vol. 50, No. 4, pp. 77\u201386, July\u2013August 2014. Original article submitted December 3, 2013. DOI 10.1007/s10778-014-0651-9 Since the IS does not have elements capable of accumulating potential energy, the stability of steady-state motions can be ascertained by analyzing the axial moment of inertia J z G of the IS about the z G -axis [1, 4]", + " To study this type of instability, it is sufficient to leave the two generalized coordinates and , which define the configuration of the system during the steady-state motions. Note also that this type of instability does not involve all possible steady-state motions. However, it is possible to show that the axial moment of inertia J z G during the other motions is less than the maximum moment J z G during the motions considered below. Let the axial moment of inertia J z G be a function of the coordinates and . The coordinates of the masses of the pendulums in O (Fig. 1a) are m l cos , m l sin , m b . (1.1) The coordinates of the center of mass of the system in O are G 0, G m m M ml M sin , G m m M mb M (M M m ). (1.2) The inertia tensor, the axial and centrifugal moments of inertia of the system are J O J J J J J 0 0 0 0 , J A m A m l b m m ( ) ( sin ) 2 2 2 2 2 , J B m B m l b m m ( ) ( cos ) 2 2 2 2 2 2 , J C m C ml m m ( ) 2 2 2 , J J 0, J m mlb m m sin (1.3) about the axes of \u00ce and J J G O G G G G G G G G G G G G G M ( ) ( ) 2 2 2 2 G G G G G ( ) 2 2 , J A mMl MG G 2 2 sin , J B ml G G 2 2 sin , J C m l MG G 2 2 2 sin , J J G G G G 0, J mMlb MG G sin (1.4) about the central axes of G G G G parallel to the of axes of O , where A A mMb M G 2 / , B B ml mMb M G 2 2 / , C C ml G 2 (1.5) are the principal central axial moments of inertia of the system during the primary motion. When projected onto the axes of G G G G , the unit vector z G directed along the z G -axis has components z G ( , sin ,cos )0 T (Fig. 1c). Then the axial moment of inertia about the z G -axis is J C C B ml u mlMb M u z G G G G G G G z J z T ( ) cos sin 2 2 1 2 2 2 m l M u 2 2 2 1 2 2 cos C C B ml u mlb u m l M u G G G ( ) cos sin cos 2 2 2 2 2 1 2 2 2 1 2 2 (1.6) ( sin , [ , ]u u 1 1, b b b mb M Mb M G / / ) , (1.7) where b is the distance from the center of mass of the system to the plane of the pendulums. We have the following dimensionless equality: ~ ( ~ ~ ) cos ~ ~ sin ~ ~ J J C B mu mbu m J z z G G G G 1 1 1 2 2 2 2 2 1 2 2 2 cos u (1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001416_0954410015603076-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001416_0954410015603076-Figure2-1.png", + "caption": "Figure 2. Rotor forces and coordinate systems.", + "texts": [ + " The operating principle of a MSDG\u2013 CMG is that the high-speed magnetically suspended rotor (MSR) system supplies constant angular momentum and the gimbal rate-servo system rotation changes the direction of the angular momentum to output gyro torque. The rotation of the inner gimbal influences the operation of the rotor. The outer gimbal is coupled with the inner gimbal; thus, high-precision control for the MSDG\u2013CMG is intractable because of the rotation of the inner and outer gimbals. The rotor force and the coordinate systems of an MSDG\u2013CMG are shown in Figure 2 and their definitions are listed in Table 2. B21 \u00bc Izzq\u0302C a l \u00f0 \u00debspan \u00fe Ixzq\u0302C a n \u00f0 \u00debspan =I 0 Izzq\u0302C r l \u00f0 \u00debspan \u00fe Ixzq\u0302C r n \u00f0 \u00debspan =I 0 q\u0302C em\u00f0 \u00decchord=Iyy 0 Ixzq\u0302C a l \u00f0 \u00debspan \u00fe Ixxq\u0302C a n \u00f0 \u00debspan =I 0 Ixzq\u0302C r l \u00f0 \u00debspan \u00fe Ixxq\u0302C r n \u00f0 \u00debspan =I 2 64 . . . 0 IxzbspanT=I Tcchord=Iyy 0 0 IxxbspanT=I 3 75 Description X-axis Y-axis Magnetic force fx, fax, fbx fy, fay, fby Magnetic torque px py Rotor angular displacements Linear displacements of rotor x y at Kungl Tekniska Hogskolan / Royal Institute of Technology on September 13, 2015pig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000667_iros.2014.6942663-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000667_iros.2014.6942663-Figure3-1.png", + "caption": "Fig. 3. A schematic diagram for robotic handwriting. p1, p2, p3 are fingertip contact points, and pr is a contact point with an internal link. z\u0302o is along the main axis of the object, x\u0302o is parallel to the shortest line from z\u0302o to pr .", + "texts": [ + " wobj is an object wrench composed of 3- dimensional force and 3-dimensional moment with respect to the global frame, and ff = [fT1 fT2 fT3 ]T is a augmented 9\u00d71 vector. The wrench matrix depends on the relative position from the object frame origin to the corresponding contact point as Gi = [ I3 [poi]\u00d7 ] , (3) where poi = pi \u2212 po, and [ \u00b7 ]\u00d7 stands for a skewsymmetric matrix for a 3-dimensional vector. Again, Gf = [Gf,1 Gf,2 Gf,3] is a augmented 6\u00d79 matrix. If the contact points and the coordinate systems are assigned as shown in Fig. 3 for clear representation, we can describe the reaction wrench from the internal link contact as follows. Grf\u0303r = [ I [por]\u00d7 ] f\u0303r = [ \u2212\u03bex\u0302o \u2212\u03be\u2016por\u2016y\u0302o ] , (4) where \u03be = \u2016f\u0303r\u2016 \u2265 0. Note that the magnitude of the reaction force, \u03be, is unlimited unless the link mechanically breaks down, and the corresponding reaction wrench, Grf\u0302r, is in a fixed direction with respect to the object frame. In order to find the finger force that satisfies (2) for a given desired wrench, wd, it is preferable to find one that uses less energy, in other words, with smaller magnitude", + " In that case, the finger force needs to be considerably large to produce a certain mount of moment about the center of the fingertip contact points. Therefore, the optimization policy for reaction force in this paper is to eliminate the moment for finger force about the fingertip centroid, as long as possible:[ 0 y\u0302o ]T G (pc) f ff = 0, (5) where G (pc) f is a wrench matrix about pc, the centroid of the fingertip contact points. Note that the reaction force can provide only one directional moment which is in negative-y\u0302o direction under the object frame convention as in Fig. 3. Equation (2) and (5) yield that the magnitude of the reaction force satisfies the following to minimize the finger force for a desired wrench. \u03be\u2217 = max ( 0, \u2016pe \u2212 pc\u2016 \u2016pr \u2212 pc\u2016 x\u0302To Uwd ) \u2265 0, (6) where pc = 1 nf nf\u2211 i=1 pi for nf = 3, and U = [I3 O3]. As a result, a reduced desired wrench is obtained by w\u0303d , wd \u2212Grf\u0303r = Gfff (7) = wd (x\u0302To Uwd \u2264 0)( I +Gr \u2016pe \u2212 pc\u2016 \u2016pr \u2212 pc\u2016 x\u0302ox\u0302 T o U ) wd (x\u0302To Uwd > 0) . (8) This equation shows that the reactional internal contact comes into effect when the desired wrench is in the opposite direction to the reaction force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000062_aespc44649.2018.9033257-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000062_aespc44649.2018.9033257-Figure1-1.png", + "caption": "Fig. 1: Schematic diagram of TRMS", + "texts": [ + " In this paper, a comparative analysis is done to check the performance of both the filters (KF and EKF) in both simulation and real time. This paper is organized as follows: Section II describes the mathematical modeling of the TRMS, Section III describe the Kalman filter based estimation technique, Section IV describe the extended Kalman filter based estimation technique, simulation and realtime result analysis has been presented in Section V. Finally, conclusion is given in the Section VI. A schematic diagram of the TRMS is shown in the Fig.1. The different parts and motions of the plant can be visualized from this schematic diagram. The rotor which is generating vertical thrust is called the main rotor. Similarly, the rotor generating horizontal thrust is called the tail rotor. The nonlinear dynamics of the TRMS can be represented as follows: d dt \u03c8 = \u03c8\u0307, (1) d dt \u03c8\u0307 = a1 I1 \u03c421 + b1 I1 \u03c41 \u2212 Mg I1 sin\u03c8 + 0.0326 2I1 sin(2\u03c8)\u03d5\u03072 \u2212 B1\u03c8 I1 \u03c8\u0307 \u2212 kgy I1 a1 cos(\u03c8)\u03d5\u0307\u03c41, (2) d dt \u03d5 = \u03d5\u0307, (3) 978-1-5386-8333-0/18/$31.00 c\u00a92018 IEEE d dt \u03d5\u0307 = a2 I2 \u03c422 + b2 I2 \u03c42 \u2212 B1\u03d5 I2 \u03d5\u0307\u2212 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001401_12.2041268-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001401_12.2041268-Figure2-1.png", + "caption": "Fig. 2: NIR-camera sensor module mounted to a TRUMPF BEO fixed lens welding optics (left) or to a TRUMPF PFO-3D scanner optic (right).", + "texts": [ + " at the melting point of iron (1,808 K) and below. InGaAs detectors are much better suited to detect this radiation than Si-based sensors. 2. SENSOR PRINCIPLE AND DESIGN The camera is mounted to the welding optics and uses a dichroitic mirror for observation of the weld zone coaxially to the processing beam. Frequently the observation port is equipped with a Si-camera for the visible spectrum to allow workpiece monitoring and measurement. The InGaAs-camera can be mounted instead of or additionally to a Si-camera (fig. 2, left); if both are mounted a wavelength-selective beam splitter is used to separate the two images. Coaxial mounting is preferred as it minimizes the influence on the interfering contour of the welding head, makes adaption to welding machines simple, and virtually does not affect the high-power processing beam. The sensor camera is connected to an image processing computer for image capture, visualization and storage, and evaluation by automated image processing. The InGaAs-camera we are using is capable to capture 100 full frames per second at a resolution of 256x320 pixels", + " One simple measure to quantify the cool-down process is the decay length of the emitted thermal radiation along the center line of the solidified weld seam (assuming an exponential decrease of the radiation intensity) 7 . If the fusion to the lower plate is missing, there is a reduced heat flow to the bottom plate and a prolonged thermal trace on the top sheet: the decay length increases. Fig. 12 shows two examples of c-cramps on lap joints which are welded by a scanner optics TRUMPF PFO-3D (see fig. 2, right). Frequently there is an intended or non-intended gap between the two sheets to be welded, but if the gap becomes too big there is no fusion between the two sheets anymore. To make things worse this missing fusion cannot be traced by visual inspection of the upper and lower bead (therefore this fault is frequently called \u201cfalse friend\u201d) but needs other types of inspection in a subsequent process step, often in a separate workstation. The evaluation of the NIR-images that are recorded during the welding process itself reveals a clear identification of a \u201cfalse friend\u201d by means of an increased decay length (fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000196_tmag.2011.2153410-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000196_tmag.2011.2153410-Figure5-1.png", + "caption": "Fig. 5. Induction motor. (a) Analyzed model. (b) Mesh for the analysis.", + "texts": [ + " Therefore, we applied the simplified SD-EEC method to the magnetic vector potentials in two approaches. One is to apply the simplified SD-EEC method to the magnetic vector potentials only in the stator every half cycle period of the power frequency as well as it was applied to an IPM motor. The other is to apply the simplified SD-EEC method to the magnetic vector potentials in the stator every half cycle period of the power frequency and those in the rotor every half cycle period of the slip frequency. Fig. 5 shows an induction motor [12] driven by the sinusoidal voltage source. Table III shows the analysis conditions. Tables IV and V show the electrical angles when the simplified SD-EEC method is applied to the magnetic vector potentials in the stator and the rotor, respectively. Fig. 6 shows the calculated primary current waveform, which are normalized by the RMS value of the primary current at the steady state. Using the simplified SD-EEC method only in the stator, the convergence of the calculated primary current waveform is improved at any slip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002344_s00170-016-9132-0-Figure27-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002344_s00170-016-9132-0-Figure27-1.png", + "caption": "Fig. 27 Model of the infrared staking machine", + "texts": [ + " Figure 26 describes the history of the MIGA-based optimization process. The maximizing testing force (F) of 352.45 N was obtained at a heating time (th) of 14 (s), a cooling time (tc) of 14 (s), and an airflow rate (Af) of 60 ft3/h, respectively (Table 6). To carry out physical experiments, a prototype machine was designed based on a concept of the infrared staking process. The machine model had five primary components: upper and lower fixtures, a frame, a halogen lamp power supplier, and a hydraulic cylinder; these and other auxiliary equipment are shown in Fig. 27. The upper fixture was equipped with infrared staking modules, whereas the lower fixture was used to clamp the car door trim. During the staking process, the lower fixture and its attached car door trim were first moved horizontally. The staking modules then were moved down vertically to the stud positions with the support of the upper fixture and hydraulic cylinder, and all steps were carried out to generate IS joints. This infrared staking system was fabricated based on the designed model (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001069_cdc.2014.7039722-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001069_cdc.2014.7039722-Figure1-1.png", + "caption": "Figure 1. Reference frames.", + "texts": [ + " Let FG be an inertial reference frame, with arbitrarily chosen origin and Euclidean coordinate system. Let FO be a moving reference frame attached to the tracked target with arbitrarily chosen origin and Euclidean coordinate system. Let FC be a reference frame attached to the moving camera, with right handed coordinate system origin at the principle point of the camera, the \u2192 e 3 axis pointing out along the optical axis of the camera, \u2192 e 1 axis aligned with the horizontal axis of the camera, and \u2192 e 2 , \u2192 e 3 \u00d7 \u2192 e 1, as shown in Fig. 1. Let the vectors \u2192 r q and \u2192 r c represent the position of a feature point on the target and the principle point of the camera with respect to the origin of the ground frame. The relative position of the feature point with respect to the camera origin is \u2192 r q/c = \u2192 r q \u2212 \u2192 r c. (1) The relative velocity as viewed by an observer in the ground frame, denoted \u2192 v q/c, is \u2192 v q/c = d dt G \u2192 r q/c = d dt G \u2192 r q \u2212 d dt G \u2192 r c. (2) To facilitate the subsequent analysis, it is beneficial to relate the rate of change of the relative position as viewed by an observer fixed in the ground frame to the rate of change as viewed by an observer fixed in the camera frame as d dt G \u2192 r q/c = d dt C \u2192 r q/c + \u2192 \u03c9 G C \u00d7\u2192r q/c, (3) where \u2192 \u03c9 G C is the angular velocity of the camera frame with respect to the ground frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003261_tii.2020.3004389-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003261_tii.2020.3004389-Figure1-1.png", + "caption": "Fig. 1 Prototype of the vehicle PMSM.", + "texts": [ + " The prediction accuracy in tracking the thermal states of the electric machines has been improved by nearly 60% in a standard test cycle. The predicted deviation fluctuates between -1%~4.5% when the step-size of the prediction is set to be 0.5s, which is acceptable for the long-range pre-adjustment of the vehicle thermal management. Besides, the computation cost is also evaluated in the hardware. It shows the proposed model can fully meet the needs of fast calculation and prediction. To form an intuitive description, our study focuses on a common water-jacket-cooled PMSM used for electric vehicle propulsion, as shown in Fig. 1. The stator uses 10-pole and 60- slot topology with single-layer windings. There are 8 turns in one slot and each turn consists of 15 round wires of 0.95 mm in diameter. The copper fill factor is approximately 53.2%. The permanent magnet array uses Nd-Fe-B with axially segmented pole pieces. A layer of carbon fiber sleeve is used to strengthen the rotor at high-speed. The basic configuration and the output capacity are listed in TABLE I. It is worth noting that the work is also applicable to other formats of electric machines", + " Then most of the heat flow will transfer from the silicon iron to the housing. The lumped thermal resistance can be derived from the reading of the thermocouples inserted in the contact interfaces of housing and the temperature of water flow at the thermal balance state. The temperature distribution in the winding is very sensitive to its thermal conductivity [2]. In most cases, it\u2019s not possible or desirable to model each individual wire. The lumped thermal conductivity is usually used to avoid excessive modeling [10, 25]. Based on the coordinate system shown in Fig. 1, the lumped thermal conductivities in the axial direction and the transverse plane can be estimated by (9) and (10) [26, 27]. -1 1 1 1 a c ci ii c ci ii v v v = + + (9) ( ) ( ) ( ) ( ) 1 1 1 1 c c c p e p c c c p v v v v + + \u2212 = \u2212 + + (10) ci ci ii ii p ci ii v v v v + = + (11) The subscripts c, ci and ii represent the copper, the surface coating insulation and the impregnated insulation compound respectively. , , c ci ii v v v are the volume ratios, which can be deduced from the packing factor PF of the windings, as 2 2 2 2 2 , , 1c c c ci ii r r r v PF v PF v PF r r \u2212 = = = \u2212 (12) in which, r is the wire radius and rc is the radius of bare wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure11-1.png", + "caption": "Fig. 11 Study of the human wrist movements for an anthropomorphic design (Masia et al. 2009)", + "texts": [ + " For that reason, the joint that corresponds to the cradle rotates in a contrary way to the rest of the joints. Another 32 R. Cab\u00e1s Ormaechea objective of the cradle is to help the driving system of the RL1 to reach all the positions and tasks needed by rotating the energy transformer in just one way. This allows the hand to pass from a standby state to and operative one, ready to hold an object through different states transforming mechanisms. It is possible to see an already developed RL1 Hand in Fig. 11. This group of mechanisms make the RL1 Hand driving system to be formed by a unique Multi-State Actuator that allows the Hand to pass through different positions, since the resting states inside the anchor until the fingers and thumb driving position, necessary to hold an object. It is also possible to see those different positions. Finally, Fig. 12 shows the RL1 materializing holding tasks for different objects with different shapes and sizes. The RL2 Hand has been designed to be mounted on a new version of the MATS Robot, The ASIBOT Robot Arm that differs from the previous version by small changes at the control system, sensor system and the capacity computing resources, without changing the mechanical structure", + " Typically, a high-ratio gearbox is used for vertical movements and a more direct drive for the horizontal movement. Additionally, decoupling reduces power losses since actuators that move do not withstand the weight and vice versa (Kar 2003). Besides the more suitable actuator selection, a decoupled configuration simplifies the inverse kinematics calculation that in turn, simplifies the trajectory control, especially during the support phase, and leads to more continuous and efficient trajectories to be described by the actuators. As shown in Fig. 11 decoupled configuration has been traditionally achieved by means of SCARA configuration and two-dimensional pantographs. Their sequences or joints are respectively rotational-rotational-prismatic and rotationalprismatic-prismatic. The structure of Fig. 11b assembles a mammal configuration. A preferred, but not necessarily better, configuration mimics the insect or reptile structure and movement, operating the rotational joint around a vertical axis. Climbing robots like the ROWER, the REST, or the tethered ROBOCLIMBER use a SCARA configuration. Many of the Hirose\u2019s robots TITAN, the tethered climber DANTE II or the MECANT use a two-dimensional pantograph. The hybrid Wheel leg of Guccione and Muscato (2003) includes a similar performance (rotational-prismatic-prismatic) although without using a pantograph", + " The main features of this arrangement are: 1) good rigidity of the structure; 2) direct drive of the manipulandum, which eliminates any backlash in the force/ motion transmission; 3) minimization of the overall inertia, because most of the Fig. 8 Robotic hand with three-one DOF articulated fingers (Nava Rodr\u00edguez et al. 2004) Mechanical Design Thinking of Control Architecture 99 100 N. E. Nava Rodr\u00edguez mass is either fixed, or close to the rotation axes; 4) independence of each single DOF (Masia et al. 2009). The ranges of motion for each DOF have been fixed based on the human wrist capabilities. Figure 11 shows plots of human wrist movements, in which ranges of motions can be recognized. Thus, the mechanical design of the wrist robot of Fig. 10 has been constrained to move from -70 to 70 for flexion/extension, from -35 to 35 for abduction/adduction and from -80 to 80 for pronation/supination. A suitable adaptive control has been implemented in the robot wrist in order to maximize task complexity as a function of the level of performance. It induces the patient to maximize the ability to face complex tasks while minimizing the reliance on robot assistance (Masia et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003693_icpre51194.2020.9233171-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003693_icpre51194.2020.9233171-Figure1-1.png", + "caption": "Figure 1. Diagram of longitudinal force analysis of vehicle on a slop.", + "texts": [], + "surrounding_texts": [ + "When the car is driving on a ramp, regardless of the impact of the lateral force of the vehicle, the vehicle dynamics equation is established as follows: = + + +t i f w jF F F F F (1) Where tF , iF , fF , wF and jF are vehicle drive force, grade resistance, rolling resistance, air resistance, and acceleration resistance respectively. In the above equation, they can be calculated from the following expressions. 342 2020 The 5th International Conference on Power and Renewable Energy 978-1-7281-9026-6/20/$31.00 \u00a92020 IEEE Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 16,2021 at 22:27:41 UTC from IEEE Xplore. Restrictions apply. 0e t T i F r (2) siniF mg (3) cosfF mg f (4) 2 D 21.15 w C Av F (5) j dv F m dt (6) where eT : Drive torque for the motor 0i : Drivetrain speed ratio : Transmission system efficiency r : Wheel Radius m : Vehicle mass g : Acceleration of gravity : Road grade angle f : Rolling resistance coefficient A : Windward area DC : Wind resistance coefficient v : Current speed : Vehicle rotation mass conversion coefficient dv dt : The current acceleration of the vehicle Substituting equations (2), (3), (4), (5), and (6) into (1), the longitudinal motion of the vehicle can be expressed as the following: 2 0 sin cos 21.15 e D aT i C Au du mg mg f m r dt (7) Considering the grade of the road is generally relatively small, so sin tan i , and cos 1 . Where 2 0 21.15 e D aT i C Au du mg i mg f m r dt (8) III. MASS ESTIMATION BASED ON ADAPTIVE" + ] + }, + { + "image_filename": "designv11_34_0001482_1.4029187-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001482_1.4029187-Figure1-1.png", + "caption": "Fig. 1 Stephenson type II six-bar linkage and one equivalent linkage", + "texts": [ + ", the reference link, the input link, the couple link, and the output link are selected, the equivalent linkage is determined by the four instant centers between the two neighboring links. These four links can be any four links from the original linkage. For type I equivalent linkage, the input link is from the four-bar loop while the input link of the type II equivalent linkage can be also from other loop. The proposed method to find the equivalent four-bar linkages is illustrated with Example 1. Example 1. Equivalent four-bar linkages for Stephenson II sixbar linkage. The Stephenson type II six-bar linkage in Fig. 1, which consists of dyads and triads, can be replaced by a four-bar linkage A0ECC0 as the equivalent linkage [13]. E is the intersection of binary link AD and B0B. It is also the instant center I25, which is link 5 relative to link 2. The couple link EC in the equivalent linkage is in the same velocity state as that of the ternary link BCD of the actual linkage. Both of them can be regarded as being rigidly connected to each other. With link 2 as the input and link 1 as the reference link, or joint I12 as the input joint, A0E and A0AB0 can be considered as a rigid unit in itself", + " It also hints that as long as the three passive joints of the type II equivalent four-bar linkage are collinear, the linkage is at dead center positions. The proposed method to identify the dead center position of single-DOF complex planar linkage can be easily used to construct the geometric configurations for the dead center position of single-DOF complex linkages. It is illustrated with the following examples. Example 2. A dead center position for Stephenson Type II linkage. According to the above criteria, with joint A0 as the input, the six-bar linkage of Fig. 1 is at the dead center position when the three joints E (I25), C (I56), and C0 (I16) of the equivalent four-bar linkage A0ECC0 lie on a common straight line, where E is the instant center I25. This also means the three lines AD, B0B, and C0C intersect at a common point E. A dead center configuration is shown in Fig. 3. For all dead center configurations of Stephenson type II linkage with the same input, the condition of three passive joints of the equivalent linkage must lie on a common straight line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000443_imece2014-38924-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000443_imece2014-38924-Figure4-1.png", + "caption": "FIGURE 4. Virtual prototype.", + "texts": [ + " \u2022 PLC application: the PLC is the component that manage signals coming from sensors and signals going to actuators. Additionally, the PLC is one of three technological alternatives to perform closed-loop control in the tank system. Hence, a control application was developed. \u2022 LabVIEW interface: this interface was required to communicate the laptop computer to the PLC. Additionally, several control algorithms can be developed and implemented over this platform. \u2022 Data management: the software interface allows the user to export and save data for further analysis. Figure 4 shows the virtual prototype of the system, obtained in the detailed design stage. The tank system was manufactured using industrial components that were bought to local providers, and local shops were used for parts fabrication. The assembly was carried out at one of the University\u2019s mechanical shops. Figure 5 shows the assembly progress of the developed system. As described in the Conceptual Design section, the equipment offers several possible configuration options for the process\u2019 dynamics. Figure 6 describes the closed-loop interaction of the controller, the control valve, and the tanks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002990_012029-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002990_012029-Figure3-1.png", + "caption": "Figure 3. Spur gears transmission [2].", + "texts": [ + " With respect to [2], the following expression for the loaded transmission error is used the transmission error amplitude. Tangent stiffness matrix and vector of inner forces are defined as follows . Also normal contact force can be calculated from equation Mesh damping can be taken into account as and gear inertias of wheel \u201cA\u201d and \u201cB\u201d respectively, the mesh damping ratio. 5. Test case This test [2] concerns the analysis of a transmission composed by two axes linked through a pair formed by two equal spur gears (figure 3). The system is driven at a constant speed of 12000 RPM at wheel 1, and transmits a torque applied at wheel 2. Gear properties are: normal modulus , pitch diameter , pressure angle and teeth number (cone and helix angles are trivially ). Their mass and rotary inertia are: and . Mesh stiffness is , gear width , mesh damping , and the amplitude of the loaded transmission error is . We have considered only the torsional stiffness of shafts in the model, with values: , , and damping: , . Disks at extremes have rotary inertias and " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003519_ccc50068.2020.9188831-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003519_ccc50068.2020.9188831-Figure2-1.png", + "caption": "Fig. 2: The coordinate system ofthe slung load", + "texts": [ + " In this section, the position and the velocity of the quadrotor in the inertial coordinate system are represented by P = [x, y, z f e K3 and V = [vx, vy, vz f e M3. The Euler angle is represented by $=[p,9, y / f e M3 . m is the mass of the quadrotor, g is the acceleration of the gravity and I = diag \\Ix,Iy,Iz} is the inertia tensor matrix. Based on the above assumptions and the definition of the coordinate system, considering environmental disturbance, a realistic dynamics model of the quadrotor is established as following: In Fig. 2, X FY and X represent the position of the load in Z I . a is the load swing angle in the X IOZI plane and is 0\u00b0 when the load in the YIOZI plane. Similarly, P is the load swing angle in the YIOZI plane and is 0\u00b0 when the load in the X IOZI plane. Because the load is not allowed to swing to the upper hemisphere above the quadrotor, a and P arerequiredbetween \u00b190\u00b0. According to the coupling system, the load swing exerts a constantly changing pulling force to the quadrotor. Since the load is suspended at the COM of the quadrotor, the load swing only affects the acceleration of the quadrotor in the x , y and z directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003638_ies50839.2020.9231564-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003638_ies50839.2020.9231564-Figure1-1.png", + "caption": "Fig. 1 Force distribution and X configuration", + "texts": [ + " The challenge in controlling a quadcopter is that the quadcopter has six degrees of freedom but there are only four control inputs and therefore it is considered as an underactuated system. The embedded controller build in this work is based on STM32F103C8T6 microcontroller. A PID controller is designed for stabilizing the attitude and altitude position. The trajectory is designed as set points of altitude control to make it can take-off autonomously. In this research, configuration quadcopter is selected instead of configuration due to its advantage in better force distribution [13], as shown in Fig. 1. 978-1-7281-9530-8/20/$31.00 \u00a92020 IEEE 233 Authorized licensed use limited to: UNIVERSITY OF BATH. Downloaded on November 02,2020 at 08:29:37 UTC from IEEE Xplore. Restrictions apply. II. SYSTEM DESIGN In this work, a preliminary design of quadcopter Embedded Flight Controller Unit (EFCU) including a conventional PID for stabilization is presented. The EFCU controls the quadcopter in hovering/altitude hold mode and automatic take-off and landing. The ultrasonic sensor HCSR04 is used for measuring the altitude of quadcopter while hovering", + " where: ; ; is cut-off frequency. C. Altitude Sensor Altitude of quadcopter to ground can be obtained using ultrasonic sensor HC-SR04. This sensor works as follows : transmitter (Tx) sends ultrasonic wave in velocity v, then travelling time (in second) of ultrasonic wave from transmitter to receiver (Rx) is measured to know altitude of quadcopter. It is formulated as follow: Altitude= (Time tracking * 0.5) * 340 Quadcopter motion is produced by controlling the speed of all 4 propellers ( and ) that produce forces as shown in Fig. 1. Speed of each propeller is controlled by controlling the value of PWM that fed to each ESC (electronic speed controller). For stabilzing the quadcopter and automatic take-off controller, the cascade PID controller is applied. The control structure is shown in Fig. 7. The formula for each PID controller is as follow: (6) is corresponding error, , , and are control parameters for proposional, integral and diferensial control, respectively. Furthermore, the output of each PID controller is coverted into PWM for each ESC by using tranformation matrix as follow: (7) where k is positive constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000014_iros40897.2019.8967565-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000014_iros40897.2019.8967565-Figure1-1.png", + "caption": "Figure 1. Passive biped model.", + "texts": [ + " The impacts introduced by both methods in reality are not present in human walking, limiting the generalization of the studies\u2019 results. To the best of our knowledge, no method of incorporating the exact kinematics of the foot-to-ground variable-curvature rolling contact in a passive dynamic biped walker has been published to date. In this paper, a passive walking model incorporating elastic and damping elements along its legs as well as circular feet, firstly introduced in [7], is extended to investigate the passive mechanisms induced by semielliptical foot shapes, see Fig. 1. 978-1-7281-4004-9/19/$31.00 \u00a92019 IEEE 6302 Authorized licensed use limited to: Western Sydney University. Downloaded on July 26,2020 at 08:24:08 UTC from IEEE Xplore. Restrictions apply. The variable-curvature elliptic foot shape provides a certain curvature when the foot is \u201cflat\u201d on the ground, but a different curvature at the beginning and end of the footground contact, emulating the rolling radius progression recorded in human gait. This model is studied for its passive behavior, to investigate the effects of the variable-curvature foot shapes on the energetics of walking", + " The biped model is presented in Section II, where the focus is turned towards the analytical modeling of the semielliptical foot shape and its interface with the ground, as well as the integration of the foot\u2019s kinematics with the rest of the biped\u2019s dynamics. In Section III, the developed model is studied for its passive behavior, the notion of energetic efficiency in stable passive walking is discussed and re-defined, and useful guidelines are drawn regarding the relationship between foot shape and energetic efficiency. Finally, Section IV presents the conclusions and sets the directions for future work. A passive biped model, as is shown in Fig. 1, is employed. The biped performs passive walking in the x-direction of the x-y coordinate system (CS) of Fig. 1, which has negative slope a. The model consists of two elastic legs of initial length Lnat: the legs\u2019 elastic and damping constants are k and b respectively. The biped\u2019s body mass M is located at the hip joint, and the legs\u2019 inertial properties are introduced through the leg mass m, located at a distance l from the bottom of each foot. The biped is situated inside a gravity field of acceleration g in the -Y direction of the global X-Y CS. The xE-yE coordinate system is a local CS attached to the leg in contact with the ground", + " For a biped with circular feet of radius r, the motion of each foot rolling on the ground is described by: (3) where xr is the foot center\u2019s displacement in the x-direction. To mimic human anatomy and to study the effects of varying foot curvature on passive walking, in this work the biped\u2019s feet are assumed to have a semielliptical shape. This change in foot shape complicates the analytical description of the rolling motion of the feet on the ground. Fig. 2 presents the elliptical foot shape on the foot-bound xE-yE CS, see Fig. 1. An ellipse in this plane is made up of points of the form: (4) where ra and rb are the ellipse\u2019s major and minor radii on the xE and yE axes respectively, and \u03c6c is a parameter visualized as the angle of a line whose intersections with two circles of radius ra and rb result in the (xi, yi) coordinates of each ellipse point, as is shown in Fig. 2. This point on the ellipse determines the geometric angle \u03c6e, for which: (5) This last equation links a geometric point (xi, yi) of the ellipse\u2019s perimeter to the parameter \u03c6c used in the definition of the ellipse, enabling the expression of contact geometry through simple algebraic equations", + " The DSP process can also be described by the discrete function f2 in a manner similar to (17): (23) G. Transition between steps The state at the beginning of the (n+1)th step, xn+1, emerges from the multiplication of xn,TO with the transformation matrix T8x8: (24) which flips the stance and swing leg initial conditions to prepare for the next step. The elements of T are all zero except for t15=t26=t51=t62=-1 and t37=t48=t73=t84=1. The negative elements of T are due to the convention of defining the leg angles in opposite directions, as seen in Fig. 1. H. Fixed points and gait stability The process described by (17), (23) and (24) can be summarized by (25): (25) where P is a discrete function describing a full step of the biped. A fixed point of P, x*, constitutes a fully repetitive gait and is defined by (26): (26) A Newton-Raphson numerical method is used to find such gaits: (27) In (27) \u03bd is a relaxation parameter, used to regulate the solution steps of the solver as P is highly nonlinear: setting v=1 as in the original method, might prohibit the convergence of the solver towards the solution, since the Jacobian \u2207P can obtain large values", + " The energetic losses of the biped due to its intrinsic damping elements as well as due to the impacts with the ground are compensated by the potential energy obtained by its descent through the gravity field. Fig. 5 presents the energy distribution for a biped walking on a negative slope. It can be seen that the kinetic as well as the elastic energy of the biped at each instant are oscillating around a fixed average value during the passive walk, indicating the repetitiveness of the gait. The gravitational potential energy presents a decreasing trend as the biped descends towards the negative Y-direction of Fig. 1, while the total energy lost through damping presents an increasing trend as the dissipation power due to the dampers and ground impacts is accumulated. The sum of the above energetic components, plotted in the last plot of Fig. 5, is shown to be constant during each step as a direct consequence of energy conservation. The sum only decreases with a discontinuity at each HS instance: this is a consequence of the conservation of angular momentum of the system before and after the HS event. The discontinuity step of the plot is then equal to the amount of energy lost due to the impact of the foot with the ground at HS. The Cost of Transport (COT) of any walking machine is defined as the ratio of the energy loss during a walk over the product of the machine\u2019s weight with the distance it travelled: (30) where Eout is the energy loss due to damping and ground impacts, and \u0394x is a distance traveled in the x-direction of Fig. 1. However, for the gait to be repetitive, the total energy must be conserved, and therefore the energy output must equal the energy input, Ein. In passive walking, Ein originates from the energetic gain due to the descent , therefore: (31) The convergence of the COT as defined in (30) towards the theoretical value of (31) for a passive biped commencing its gait near a stable fixed point is shown in Fig. 6. It can be observed that as the gait progresses, the COT converges to (for ), as a direct consequence of the energetic equilibrium in the stable system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001852_0954410016638870-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001852_0954410016638870-Figure1-1.png", + "caption": "Figure 1. Missile interceptor with aerodynamic surfaces and reaction jets.", + "texts": [ + " In \u2018\u2018Control allocation\u2019\u2019 section, the L2 optimal control allocation strategy is designed for the dual controlled missile. In \u2018\u2018Virtual control law design\u2019\u2019 section, the robust sliding sector with a parameter update law is proposed and the virtual control law is designed via utilizing the proposed adaptive sliding sector control. Numerical simulation results are shown in \u2018\u2018Simulation results\u2019\u2019 section, and conclusions are made in the last section. A missile equipped with aerodynamic surfaces and reaction jets is shown in Figure 1. The reaction jets are located ahead of the center-of-gravity of the missile, which can provide additional forces and moments on the missile and make the missile more agile. In the coordinate system described in Figure 1, the positive normal acceleration Nz is the normalized aerodynamic force along the negative body-fixed z-axis. We can see that a positive force generated by reaction jets causes the missile to build up to a positive angleof-attack which will produce a positive normal acceleration. In Figure 1, the elevator deflection u is in the positive direction, steering the missile to build up a negative angle-of-attack which will produce a negative normal acceleration. The missile dynamics in the pitch plane taken into account for the control law design can be written as2 _ \u00bc QmSm mVT CN \u00fe q QmSm mVT CN u 1 mVT uT \u00bc Z \u00fe q\u00fe Z u \u00fe ZTuT _q \u00bc QmSmLm Im Cm \u00fe Cmqq\u00fe Cm u \u00fe lm Im uT \u00bcM \u00feMqq\u00feM u \u00feMTuT _ \u00bc q 8>>>>>>< >>>>>>: \u00f01\u00de where CN and CN are aerodynamic force coefficients, Cm , Cmq, and Cm are aerodynamic moment coefficients, Qm is the dynamic pressure, lm is the moment arm, Im is the moment of inertia, Sm is the reference surface, and Lm is the reference length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002220_j.ifacol.2015.09.281-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002220_j.ifacol.2015.09.281-Figure2-1.png", + "caption": "Fig. 2. Two-link robot manipulator model", + "texts": [ + " Problem of controller design for tracking of robotic manipulators is treated in numerous research works (Young (1978); Slotine and Sastry (1983); Spong and Vidyasagar (1989); Piltan and Sulaiman (2012); Chanda and Gogoi (2014)). This problem gives an appropriate example of the high nonlinear closely coupled dynamical plant and allows to demonstrate efficiency of the proposed trajectory tracking controller design methodology which is based on the singular perturbation technique. Let us consider a two-link manipulation robot model as shown on Fig. 2, where the dynamical behavior is described by the following equations (Young (1978)): x (1) 1 = x2, x (1) 2 = [a22/a][\u03b212x2(x2 + 2x4) + \u03b31g + u1 + w1] \u2212[a12/a][\u2212\u03b212x 2 4 + \u03b32g + u2 + w2], x (1) 3 = x4, (36) x (1) 4 =\u2212[a12/a][\u03b212x2(x2 + 2x4) + \u03b31g + u1 + w1] +[a11/a][\u2212\u03b212x 2 4 + \u03b32g + u2 + w2], where a11 = (m1 + m2)l21 + m2l 2 2 + 2m2l1l2 cos(x3), a12 = m2l 2 2 + m2l1l2 cos(x3), a22 = m2l 2 2, a = a11a22 \u2212 a2 12, \u03b212 = m2l1l2 sin(x3), \u03b31 = \u2212(m1 + m2)l1 cos(x1) \u2212 m2l2 cos(x1 + x3), \u03b32 = \u2212m2l2 cos(x1 + x3), x = [x1, x2, x3, x4]T = [\u03b8, \u03b8\u0307, \u03d5, \u03d5\u0307]T , y1 = x1, y2 = x3, u = [u1, u2]T , w = [w1, w2]T , y = [y1, y2]T is the measurable output, u is the vector of joint torques (control variables), w is the vector of external joint torques (disturbances), and g is the gravitational constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000034_roma46407.2018.8986730-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000034_roma46407.2018.8986730-Figure2-1.png", + "caption": "Figure 2: Sensing gloves structure", + "texts": [ + " The PCB containing slide potentiometer and the Arduino board is attached to the glove which can be worn by the human. The finger tip of the glove is attached to the varying point of slide potentiometer. If the finger is moved, it will pull the moving point of the slide potentiometer, thereby sensing the motion of the fingers. Zigbee is attached to the Arduino board which acts as a sender to send the data of slide potentiometer and accelerometer. Inorder to obtain the gesture, the slide potentiometers and the accelerometers are used which are shown in figure 2. A. Accelerometer (ADXL 335) The rate of acceleration due to gravity can be of either static that is for tilt sensing applications or dynamic that is for motion, shock or vibration applications. Both the static and dynamic rate of acceleration can be measured with the help of the accelerometer ADXL335. The accelerometer consists of three capacitors CX, CY, and CZ that are used to determine the bandwidth of the accelerometer. These three capacitors can be accessed via XOUT, YOUT, and ZOUT pins. The functional block diagram of the accelerometer is given in the below figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003969_0954406220983367-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003969_0954406220983367-Figure5-1.png", + "caption": "Figure 5. Contact sketch of linear guideway (a) sketch of linear guideway; (b) contact model of slider/guideway and roller; (c) half contact deformation model under linear load.", + "texts": [ + " The preload of slider/guideway and roller generates the contact stress, and results in uneven distribution of asperities\u2019 contact area. According to the fractal theory, if Al is set as the maximum contact area of asperity, the number distribution function of contact asperities will be expressed as 48 n A\u00f0 \u00de \u00bc DAl D=2 2A D=2\u00fe1\u00f0 \u00de (25) Actually, equation (25) is only suited to the contact analysis between rough surfaces, however ignores two contact solids shapes, such as the case of the slider or guideway/roller interface shown in Figure 5(a). The major reason is that n A\u00f0 \u00de is also influenced owing to geometric shapes and materials for contact solids and contact load. Regarding the contact between rough cylindrical and flat surfaces, namely the slider/guideway and roller in Figure 5(b), the solid surface contact coefficient s should be introduced to be written as s \u00bc AnX A Xn (26) where An is the nominal contact area, Xn is the index,X A is the sum contact area of two solid surfaces and is written as X A \u00bc A1 \u00fe A2 (27) where A1 and A2 represents the contact area of the slider/guideway and the roller by itself. According to Hertz\u2019s contact theory, half of contact width between the cylindrical and flat surfaces shown in Figure 5(c) is expressed as b \u00bc ffiffiffiffiffiffiffi 4wR pleE q (28) where 1 R \u00bc 1 R1 \u00fe 1 R2 , R denotes the equivalent curvature radius, w is the linear load in the redial direction, le is the effective contact length of roller and slider and guideway. So the rectangle contact areas of the slider/guideway and the roller can be calculated as A1 \u00bc A2 \u00bc le 2b \u00bc 4 ffiffiffiffiffiffiffi lewR pE q (29) Then, the sum area of the two solid contact surfaces X A is expressed as X A \u00bc A1 \u00fe A2 \u00bc 8 ffiffiffiffiffiffiffi lewR pE q (30) Because of the solid surface contact coefficient s as exponent of the sum contact area to be written as s \u00bc 1 8 ffiffiffiffiffiffiffiffi lewR pE q 1 8 ffiffiffiffiffiffiffi lewR pE p (31) For the contact of the flat surface and cylinder, the number distribution function of contact asperities n0 A\u00f0 \u00de is calculated as n0 A\u00f0 \u00de \u00bc s n A\u00f0 \u00de \u00bc 1 8 ffiffiffiffiffi lewR pE p 1 8 ffiffiffiffi lewR pE p DA D=2 l 2A D=2\u00fe1\u00f0 \u00de (32) Real contact area model between rough surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003855_ddcls49620.2020.9275206-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003855_ddcls49620.2020.9275206-Figure5-1.png", + "caption": "Fig. 5: The coordinate system for MDP modeling", + "texts": [ + " A MDP includes four parts: 1) a state space S; 2) an action space A; 3) a cost function of one step c (s,a): S \u00d7 A \u2192 R; 4) a stationary one-step transition probability: p(st|s1,a1 . . . st\u22121,at\u22121). The Markov property means that currant state only depends on the last state and action. The state of position control is designed as: s = [ \u2206\u03c7,v, v\u0307 ] T (14) Authorized licensed use limited to: East Carolina University. Downloaded on June 15,2021 at 19:20:41 UTC from IEEE Xplore. Restrictions apply. where \u2206\u03c7 = \u03c7\u2032 \u2212 \u03c7 and \u03c7\u2032 = [x\u2032, y\u2032, \u03c8\u2032 ] T is the description of the target point as shown in Fig. 5. x\u2032 and y\u2032 are coordinates of the target point in the system E-xy. \u03c8\u2032 is the angular displacement form x axis to a connecting line between the robot\u2019s geometric centre and the target point. The action is the control inputs of the biomimetic propellers, therefore: a = f (15) The purpose of the position control is to control the UBVMS move to and stop at any target point from a random start location and random posture, therefore the cost function in this MDP is designed as: c (\u03c7,v) = \u03c11(x\u2032\u2212x)2+\u03c12(y\u2032\u2212y)2+\u03c13(\u03c8\u2032\u2212\u03c8)2+\u03c14k(v) (16) where \u03c1i(i = 1, 2, 3, 4) is the weight of each item, \u03c11(x\u2032 \u2212 x)2 and \u03c12(y\u2032\u2212y)2 are used for minimizing the distance between the robot and the target point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000422_transducers.2015.7180945-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000422_transducers.2015.7180945-Figure4-1.png", + "caption": "Figure 4: Size of the packaged lens compared to a commercial lens.", + "texts": [ + "00 \u00a92015 IEEE 399 Transducers 2015, Anchorage, Alaska, USA, June 21-25, 2015 The fabrication is a straightforward pick-and-place process (fig. 3): First the piezo and the glass are structured by laser ablation and glued together using a rigid polyurethane. The elastomer body is molded directly onto the lens substrate in a milled polymer mold. Then, the membrane is glued on top of the body using the same elastomer, which fills the optical defects caused by the relatively rough mold. The lens is finally placed in a compact package (fig. 4) and equipped with soldering contacts. We have fabricated prototypes using 30 \u00b5m thin borosilicate glass and an approximately 120 \u00b5m thick transversely polarized PZT piezo film. This has a piezo coefficient \ud835\udc5131 \u2248 \u2212280 \u00d7 10\u221212\ud835\udc5a/\ud835\udc49 and is coated with planar silver electrodes. The lens body is 3.5mm thick and fills exactly the membrane diameter \u2013 taking also care of the alignment during the gluing. As a clear aperture, we chose 6.25 mm (1/4\u2019\u2019) with an outer piezo diameter of 7.9 mm and a packaging diameter of 9 mm, giving a clear aperture ratio greater than 2/3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001824_1.4033034-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001824_1.4033034-Figure4-1.png", + "caption": "Fig. 4. Sketch of phase portraits of the oscillating system described by different variables which are conencted by the transformations (27) and (30). For a chosen trajectory, also the amplitudes \u03c8\u0304 and \u03c8\u0304\u2217 are denoted.", + "texts": [ + " The zeroes in the main diagonals of the coefficient matrices in (28) are in correspondence with the above proved symmetry properties and the neutral stability of the trivial solution. The formula for \u03b2\u0302 can be calculated directly from (24), (27) and (28), and we get \u03b2\u0302 = 1 8 ( 3b30 b10 + a21 a01 \u2212 b12 a01 \u2212 3a03b10 a2 01 ) . (29) Note that the normal form (28) of the governing equations can be obtained as special case from the normal form of the Hopf bifurcation (see [24], p. 152 or [25], p. 385). The effect of the amplitude of the oscillation on the angular frequency can be determined if we transform (28) into polar coor- dinates (see Fig. 4). Let r = \u221a (\u03c8\u2217)2 +(\u03d1\u2217)2, tan\u03b5 = \u03d1\u2217 \u03c8\u2217 , (30) then, (28) becomes r\u0307 = 0+O(5), \u03b5\u0307 = \u03c9L(1+ \u03b2\u0302r2 +O(4)), (31) where the time derivative \u03b5\u0307 in (31) gives the nonlinear angular frequency \u03c9N of the oscillation. According to the transformations (27) and (30), the linear approximation of the yaw angle is \u03c8 \u2248 \u03c8\u2217 = r cos\u03b5, thus, the amplitude dependence of the angular frequency can be expressed by \u03c9N(\u03c8\u0304) = \u03c9L \u00b7 ( 1+ \u03b2\u0302\u03c8\u0304 2 +O(4) ) , (32) where \u03c8\u0304 \u2248 \u03c8\u0304\u2217 = r is the amplitude of the oscillation expressed by the yaw angle (see Fig. 4). In practice, the amplitude of the oscillation is usually expressed by the lateral displacement y rather than the yaw angle \u03c8 (see Fig. 5). By considering the symmetries described in the previous subsection, the lateral displacement can be expressed in the form y(\u03c8,\u03d1) = c01\u03d1+ c03\u03d1 3 + c21\u03d1\u03c8 2 +O(5), (33) where the coefficients c01,c03 and c21 are determined by (14), and the coefficient c01 gives the linear approximation of the ratio between the amplitudes of y and \u03d1. Moreover, the connection between the amplitudes of \u03c8 and \u03d1 are determined by (24), thus, we get y\u0304(\u03c8\u0304) = c01 \u221a \u2212a01 b10 \u03c8\u0304+O(3), (34) where y\u0304 is the amplitude of the oscillation expressed by the lateral displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003694_ecce44975.2020.9235404-Figure18-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003694_ecce44975.2020.9235404-Figure18-1.png", + "caption": "Fig. 18. Magnet loss distributions between continuous-skewed and 10-step-skewed magnets.", + "texts": [ + "5 1 5550 Authorized licensed use limited to: University of Canberra. Downloaded on May 23,2021 at 22:42:22 UTC from IEEE Xplore. Restrictions apply. Therefore, the flux in the skewed stator cores tends to avoid the right corners of the magnets, which leads to unevenly distributed flux density in the stator cores and magnets, and hence the higher magnet eddy current losses shown in Fig. 17. The magnet losses are reduced by using step-skewed segmented magnets instead of a whole piece of continuous-skewed magnet. Fig. 18 shows the magnet eddy current loss density distributions comparisons between a continuous-skewed magnet and the step-skewed magnets. The 10-segmented step-skewed magnet reduces the magnet loss by over 90%. In order to evaluate the effect of skew on the magnet and iron losses of the unaligned stator-rotor configurations, 0-degree, 8- degree, 18-degree, 28-degree, and 38-degree stator-rotor skew angles are modeled and calculated by 3-D FEA. Fig. 19 shows the flux density distributions of the five unaligned stator-rotor integrated machines with different stator-rotor skew angles but the same 38-degree rotor stagger angles at no-load conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003290_tmag.2020.3006272-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003290_tmag.2020.3006272-Figure3-1.png", + "caption": "Fig. 3. Magnetic characteristics and samples of (Ce, Nd)-Fe-B magnets.", + "texts": [ + " The distinctive characteristics of (Ce, Nd)-Fe-B magnet which make it suitable for the application in VFMMs can be summarized as follows: 1) The Hc of (Ce, Nd)-Fe-B magnets can be flexibly adjusted by dual-main phase method and adopting different substitution ratios of Nd and Dy elements to satisfy the design requirement of VFMMs, in terms of the tradeoff between the required demagnetizing and re-magnetizing current amplitude and limited inverter rating as well as the cost. The (Ce, Nd)Fe-B magnets\u2019 Hc spans from -130 kA/m to the similar level of NdFeB magnets, which covers the variation range of all low-Hc magnets used in current literatures. The Fig. 3 presents demagnetization curves and samples of three grades of (Ce, Nd)-Fe-B magnets. The Hc of three grades of magnets are -326 kA/m, -262 kA/m, -207 kA/m, respectively. And hence they will forth be termed as Ce326 magnet, Ce262 magnet, Ce207 magnet for short. 2) Regardless of Hc being adjusted in a wide range, the remanence (Br) can be kept relatively constant above 0.8 T. The Ce326 magnet and Ce262 magnet have identical Br of 0.83 T, while Br of Ce207 magnet is 0.8 T. The relatively high and constant Br of (Ce, Nd)-Fe-B magnet makes it possible to offer a comparable open-circuit air-gap flux density and torque density to conventional PM machines" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001308_iemdc.2013.6556236-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001308_iemdc.2013.6556236-Figure10-1.png", + "caption": "Fig. 10. Calculated permeability distributions (No load).", + "texts": [], + "surrounding_texts": [ + "A. Variation in Field Distribution by Considering Stress First, the effects of the stress on the field distribution are investigated. Figs 10 and 11 show the permeability and loss distributions calculated by the electromagnetic field analysis and the post core loss calculation, respectively. It is observed that the permeability decreases, whereas the core loss density increases at the yoke by considering the stress caused by the shrink fitting." + ] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure10-1.png", + "caption": "Fig. 10 Conventional schemes for articulated legs: a with rotary motor (Asimo). b with linear actuators (Big-Dog)", + "texts": [ + " The actuator system should provide all the necessary states the RL1 Hand might need to achieve holding tasks and, apart from that, should develop all the movements to pass from the different states: standby to operative. A specific design for the RL1 Hand of an actuator system is shown in Fig. 9. The driving system of the RL1 Hand is formed by an energy transformer that provides three states to the two mechanisms named MA1 and MA2. Both mechanisms give movement to all the robotic system elements. Only one of them works at a time, thanks to a specific mechanism. This allows the energy transformer to send its movement selectively to each Max depending on the state of the Hand. The extraction system is shown in Fig. 10. In it is possible to see, apart from all the elements that form it, a mobile platform named Main Platform, which moves in only one direction through the Guiding Columns. The central Main Axis that transmits the movement of the energy transformer and the state selector mechanism, which is solidary to the Main Platform, makes this Platform to move up and down positioning the fingers outside and inside the anchor. The states selector mechanism, thanks to a mechanical fusible hosted inside it, passes from a spindle which runs over a central Slotted Axis to a pulley that hosts the tendons of the three fingers and transmits the moment generated by the energy transformer to each joint, every time the main Fig", + " The legs of reptiles have a similar configuration to mammals, although the parallel axes (labelled as T2 and T3) belong to very separated vertical planes and hence they undergo higher torques (Guardabrazo and Gonzalez de Santos 2004). Reptile-like robots also require an articulated body to keep balance during the crawl movement (Ishihara and Kuroi 2006). Although different types of actuators can be found, most of biped robots use the scheme shown in Fig. 9b in which electrical motors drive rotational joints. As reproduced in Fig. 10a biped robot legs require a relatively wide and flat foot to provide static stability and improve the dynamic one. Unless a penguin gait is not a concern (Collins et al. 2005), Mobile Robots 49 the clearance of the foot must be accomplished by means of an ankle, with a rotational joint whose axis is perpendicular to the sagittal plane. In order to keep balance a lateral movement is required that is performed by a set of two joints per leg (not shown in the figure), one in the hip and another in the ankle, whose axes are parallel to the movement direction", + " Despite their current application as toys or research bed tests, biped robots are intended to operate as assistance robots. Since they must be in close interaction with humans, these robots must, for psychological reasons, mimic the human appearance, configuration and movements. This is because legs for biped robots have an articulated configuration like humans, disregarding factors like autonomy that is not a concern in domestic applications. A different type of movers is included in the Big-Dog that uses hydraulic linear actuators as it shown in Fig. 10b. Since it is a quadruped, the ankle answers other purposes, mainly compliance and reaction absorption. Articulated legs have an important problem related to the size of actuators. If the robot wants to raise its body, the robot must exert torques TK and TA (or forces FK and FA) in knee and ankle, that can be as high as T LLm g \u00f04\u00de For a certain value of mass m and link length LL. Due to the coupled structure, during stance, transfer, leg ascent or leg descent, the robot must operate all of the leg actuators", + " The grasp forces of several objects with different shapes and dimensions have been measured in order to compare results with experimental validations of the robotic hand of Fig. 8. The experimental results show the practical feasibility of the prototype as three-fingered robotic sensored hand with three 1 DOF anthropomorphic fingers, having human-like operation (Nava Rodr\u00edguez et al. 2004). The human-like characteristics of the robotic hand of Fig. 8 simplify its control architecture since some aspects of grasp have been checked in the mechanical design, for example performance of fingers and grasping force. The robotic device of Fig. 10 is another example of a system inspired in the study of human body. This device has been used for rehabilitation of injured human wrist. The chosen class of mechanical solutions is based on a serial structure, with direct drive by the motors: one motor drives the pronation/supination, one motor the flexion/extension and two parallely coupled motors the abduction/adduction. The main features of this arrangement are: 1) good rigidity of the structure; 2) direct drive of the manipulandum, which eliminates any backlash in the force/ motion transmission; 3) minimization of the overall inertia, because most of the Fig", + " 8 Robotic hand with three-one DOF articulated fingers (Nava Rodr\u00edguez et al. 2004) Mechanical Design Thinking of Control Architecture 99 100 N. E. Nava Rodr\u00edguez mass is either fixed, or close to the rotation axes; 4) independence of each single DOF (Masia et al. 2009). The ranges of motion for each DOF have been fixed based on the human wrist capabilities. Figure 11 shows plots of human wrist movements, in which ranges of motions can be recognized. Thus, the mechanical design of the wrist robot of Fig. 10 has been constrained to move from -70 to 70 for flexion/extension, from -35 to 35 for abduction/adduction and from -80 to 80 for pronation/supination. A suitable adaptive control has been implemented in the robot wrist in order to maximize task complexity as a function of the level of performance. It induces the patient to maximize the ability to face complex tasks while minimizing the reliance on robot assistance (Masia et al. 2009). Particular systems, for example vision or cognitive systems require to work under certain condition that do not interfere with their correct operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003739_0954405420971128-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003739_0954405420971128-Figure4-1.png", + "caption": "Figure 4. Cylinders built by SLM along the z-axis: (a) the deposition strategy for the \u2018Rooster\u2019 built cylinder, (b) photograph of the built cylinder\u2019s end, (c) photograph of the coin blank surface after wire-EDM, (d) photograph of the coin blank surface after polishing.", + "texts": [ + " The cylinders from which the coin blanks were obtained had 60mm height and were built by SLM along the z-axis in order to minimise the use of support structures and to improve the overall efficiency and quality of material deposition. In fact, building along other than the z-axis would increase the use of support structures and inhibit the fabrication of very small holes like those used in the eyes of the \u2018Rooster\u2019 and \u2018Eagle\u2019. Wire electro-discharge machining and surface polishing were subsequently utilised to slice the cylinders into individual coin blanks with appropriate quality requirements. The holes were not polished although they could have been by subjecting the built-up cylinders to abrasive flow machining. Figure 4 shows a scheme of the deposition strategy (refer to \u20181\u2019, \u20182\u2019, etc.) together with photographs after SLM, wire-EDM and polishing, in which the improvement of surface finishing is easily observed. Deposition parameters used in the fabrication of coin blanks by SLM were identical to those provided in Table 1. The coin blanks were then compressed (coin minting) between dies in a press-tool (Figure 5) to impart lettering and other reliefs on both surfaces. The tool was designed and fabricated by the authors and consists of a simple laboratory system with an upper drive plate, a lower fixed plate, an upper die holder, a pair of reverse and obverse dies and a collar with an inner flat surface (Figure 5(a))" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003667_j.mechmachtheory.2020.104151-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003667_j.mechmachtheory.2020.104151-Figure3-1.png", + "caption": "Fig. 3. Experimental setup. (A-B) Experimental devices to measure the output of nonlinear springs from two perspectives. (C) The effective linear torqueangle curve applied to the planetary gear train.", + "texts": [ + " As the tooth profile of the ring can only be produced by internal shape cutters, the meshing process of a virtual rack and the internal planet gear was numerically calculated, and then the ring was cut using the obtained planet gear for the internal meshing pair. This section establishes an experimental setup to verify the availability of the proposed design methodology. Three representative nonlinear torque-angle relationships, including a constant torque, a negative stiffness curve and a U-shaped curve, were demonstrated with increasing complexity. In the following experiments, as shown in Fig. 3 A and B, a linear torque-angle spring was realized by connecting one end of a linear tensile spring to a rope that wound around a driving shaft for the convenient of adjustment and measurement. The tensile spring\u2019s stiffness is k l = 0 . 18 N / mm , and the shaft radius is r i = 15 mm . Therefore, the effective torsional stiffness is k eff = k l r 2 i = 40 . 5 Nmm / rad . Moreover, the output torque was measured by detecting the tension of the output rope, whose one end was wound around a rotation shaft of the planetary arm with a radius r o = 22 ", + " An assembled frame was fixed to an optical platform to support the noncircular planetary gear train. The ring was bolted to the platform, and the other noncircular gears were connected to the driving shaft and the planet shaft by spline joints. The prototype of noncircular gears and rings with involute tooth profiles was made from polymethyl methacrylate by laser cutting. A shaft protruding from the rotational axis of the 3D-printed planetary arm connected the rope with the force sensor. All the rotation shafts were mounted to the frame through bearings. As shown in Fig. 3 C, in was set at 280 \u25e6 and the spring was uniformly pretensioned to 15 N in the latter examples, with a linear torque applied to the driving shaft ranging from 225 Nmm to 422 . 9 Nmm , that is, E sum = 1583 . 2 J . In contrast, the planetary gear train was supposed to output different specific torque-angle curves at out = 100 \u25e6. The effective meshing range of the sun gear and the ring was deduced to be s = 180 \u25e6 and r = 100 \u25e6 referring to Eq. (7) . p was set to be equal to in and the center distance D was set at 50 mm " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000452_icinfa.2015.7279812-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000452_icinfa.2015.7279812-Figure1-1.png", + "caption": "Fig. 1. Coordinate system definitions of the UAV in flight", + "texts": [ + " In the most general path following problems, the objective of developed method is to accurately follow a a predefined path. For a small fixed-wing UAV, only high-level navigation control is considered here and low-level attitude control is achieved via an autopilot system. The UAV dynamics with high-level control in a 2-D horizontal plane is [13] x\u0307 = V cos(\u03c8) y\u0307 = V sin(\u03c8) \u03c8\u0307 = u (1) where, (x, y) \u2208 R2 denote the position of a UAV, V is the velocity of UAV in the forward direction, \u03c8 is the yaw of UAV. The definitions of those variables shown in Fig.1. The main objective of proposed controller is existing a optimal control input u(k), which make the following achieved. |\u03b4(k)\u2212 \u03b4d| < \u03b5 umin < u(k) < umax (2) where, \u03b4(k) = \u221a (x(k)\u2212 xd(k))2 + (y(k)\u2212 yd(k))2 is distance between the UAV position and desired position, (xd(k), yd(k)) is desired position in time k. \u03b4d is desired value of distance, \u03b5 is endurable bound of error, umin and umax are the minimum and maximum of control input, and these are considered as constraints in proposed controller. The key challenge to achieve these objectives is there exist unmodeled dynamics and alterable structure of UAV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003161_j.addma.2020.101345-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003161_j.addma.2020.101345-Figure4-1.png", + "caption": "Fig. 4. Cross sections of the inlet geometries including the nozzles: a) perpendicular and b) axial.", + "texts": [ + " They have been separated from junctions (and in some cases the inlets) in order to create sample channels which only vary by length. Only a part of all channels could be made accessible for the test rig, explained below. There are different inlet geometries. This is because the priority was on preserving as many channels with maximum length. The exits into ambient air are all geometrically similar. The outlet surfaces are perpendicular to the channel flow axis. There are two variations for the inlet geometry: an axial inlet and an inlet which is perpendicular to the channel axis. These variations are depicted in Fig. 4 a) and b). The channels are rectangular wavy channels, which were inspired by [11], but have different dimensions. An example of the channels which have been separated from the junction into single channels as described above is depicted in Fig. 5. Table 1 summarizes the CAD geometries of the tested channels and defines the abbreviation used for each channel group in this work. The hydraulic diameters range from dh = 0.92 to 1.05mm. They all share the same wavelength of 2.54 mm. The build direction defines the orientation of the channel axis relative to the build platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003390_1350650120945517-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003390_1350650120945517-Figure1-1.png", + "caption": "Figure 1. (a) Noncircular two-lobe hydrodynamic journal bearing; (b) rotor-bearing system.", + "texts": [ + " The effects of lubricant couple stress parameters as well as bearings\u2019 eccentricity ratio and noncircularity index on the dynamic trajectory, critical mass parameter, and whirl frequency ratio of the analyzed rotor-bearing system are presented. Noncircular journal bearings usually consist of a combination of two or more partial journal bearings. Each of these partial bearings is called a lobe. To determine the overall performance of noncircularlobed journal bearings, the algebraic or vector summation output of each lobe can be used according to the type of the desired quantity. Figure 1 displays a bearing supporting system for the rotary shaft, including a pair of noncircular two-lobe journal bearings. The modified governing Reynolds equation for twodimensional isothermal flow of the couple stress lubricant in the journal bearings clearance space can be expressed as20 @ @ h, l @ p @ ! \u00fe R2 @ @ z h, l @ p @ z ! \u00bc 6 R U @ h @ \u00fe 12 R2 @ h @ t h, l \u00bc h3 12 h 2 l\u00fe 24 h3tanh h 2 l ! \u00f01\u00de In the above equations, h represents the thickness of the trapped couple stress oil film between the rotor and the inner surface of the bearing shell and l indicates the couple stress parameter of the lubricant, and as it approaches zero, equation (1) reduces to the classical Reynolds equation for the Newtonian fluid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000639_1350650113480290-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000639_1350650113480290-Figure2-1.png", + "caption": "Figure 2. Structure of sealing lip.", + "texts": [ + " A WR auto water pump bearing seal was taken as an object to calculate the lubrication parameters, such as film pressure, film thickness, friction force, heat dissipation, average temperature and leakage rate with different initial interference and different rotating speed. The relation between contact interference, heat dissipation and lubrication performance is studied. Figure 1 shows auto water pump bearing seal constituted by metal frame and rubber sealing body, the sealing zone, shaped as a belt, is formed by interference fit of sealing lip and shaft. Macroscopic structure simulation of lip seal Sealing is realized by interference contact of sealing lip and shaft, the axial structure of sealing lip is shown in Figure 2. Here, 1 expresses the angle of sealing lip at Mount Royal University on June 9, 2015pij.sagepub.comDownloaded from and shaft toward lubricant, whereas 2 expresses the angle of sealing lip and shaft toward air, zbmin is the position where the initial average film thickness is the minimum. Deformation angle, contact position and contact width of sealing lip vary with different interference. Contact deformation can be obtained through FEA software ANSYS. As the structure of lip seal has the axisymmetric characteristics, two-dimensional (2D) axisymmetric model can be used to simplify the 3D model, which can reduce the requirement for calculation of memory and computing time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000000_robio49542.2019.8961677-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000000_robio49542.2019.8961677-Figure4-1.png", + "caption": "Fig. 4. Configuration of AE sensors", + "texts": [ + " To evaluate the RV reducer damage level, the return difference equipment is developed, as shown in Fig. 3. In the evaluation process, the RV is fixed on the pedestal, the 1https://www.nabtescomotioncontrol.com/pdfs/RVseries.pdf corresponding torque is measured by placing the equivalent standard metal weights on plate. Through this way, the angular value could be observed by the angular transducer. In the test rig, two AE sensors are rigidly attached to the surface of reducer output outer ring, shown in Fig. 4. By means of this configuration, the sensors could sample the AE data more precisely, since the signal attenuation from the signal source to the sampling articles are considerably alleviated. In our setup, the two AE sensors are placed in a parallel manner and the sampled data are synchronized by timestamps. Apart from that, the vibration and temperature sensors are also placed on the surface of RV, the corresponding vibration measurements and instrument temperature values are collected in the acquisition equipment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003320_j.cmpb.2020.105646-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003320_j.cmpb.2020.105646-Figure4-1.png", + "caption": "Fig. 4. Definition of articulator coordinate systems.", + "texts": [ + " The intercondylar axis r the hinge axis is the line between the two condylar balls. The -axis is also set along with the hinge direction. Then we dene the left and right reference coordinate systems, { X L ,Y L ,Z L } and X R ,Y R ,Z R } whose origins are at the centers of the left and right ondylar balls when the articulator is in the initial (centric reation) position and the directions are the same as those of the orld coordinate system. The relationship between these coordiate systems is shown in Fig. 4 . As the mouth opens, the lower jaw oves relative to the upper jaw and the condylar balls are moving nside the three dimensional glenoid fossa which will be modeled a a l i p t a c l A o m m t l t t e s d 3 b B T s the left and right curved condylar guidance paths. Therefore, for n articulator in general, when we fix the lower part (because the ower part usually sits on a table), the movement of the upper part s contributed by two condylar balls moving in the curved condylar aths. Two condylar balls are connected with the hinge axis so that he movement of the upper part is guided by the condylar guidnce paths which simulates human\u2019s condylar paths" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003390_1350650120945517-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003390_1350650120945517-Figure6-1.png", + "caption": "Figure 6. Nonlinear dynamic model for tracking the perturbed motion of the rotor center in noncircular two-lobe journal-bearing space.", + "texts": [ + " The algorithm of the stability analysis of the rotor-bearing system based on the linear model is illustrated in Figure 5. In analyzing the two-lobe journal-bearing behavior using the nonlinear dynamic model, the simultaneous solution of time-dependent governing Reynolds equation (equation (2)) and rotor motion equations are performed in successive steps to determine the new position of the rotor center X, Y in the bearings\u2019 space after leaving the static equilibrium point XJ0,YJ0\u00f0 \u00de, as shown in Figure 6. The difference between the resulting forces FX,FY\u00f0 \u00de and FX0,FY0\u00f0 \u00de\u00f0 \u00de from equation (7) by applying the dynamic p\u00f0 \u00de and static p0\u00f0 \u00de pressure distributions at each solution step determine the speed _X, _Y and acceleration \u20acX, \u20acY of the center of the rotor in the direction of the coordinate axes, and by substituting them in the equations of motion, the new position of the rotor center is obtained in the next step. The boundary conditions used to determine the dynamic pressure distribution of the couple stress lubricant film in the successive steps of analysis considering the Reynolds boundary condition to determine the cavitation zone on each lobe are p i1, z, \u00bc 0, i \u00bc 1, 2 p i2, z, \u00bc @p i2, z, =@ \u00bc 0 p , 1, \u00f0 \u00de \u00bc 0 p , z, \u00f0 \u00de \u00bc p \u00fe 2 , z, \u00f0 \u00de \u00f026\u00de Based on the two-lobe journal-bearings\u2019 geometry and the characteristic of the rotor center oscillations, the dynamic pressure distribution of the oil film is symmetric and harmonic in the longitudinal and circumferential directions, respectively, and the value of lubricant film pressure at the end nodes of the problem domain is equal to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001225_dscc2014-6114-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001225_dscc2014-6114-Figure5-1.png", + "caption": "Figure 5. Tracking constraint in the original form", + "texts": [ + " Condition (16) verifies whether the system can be driven from all vertices of the set X\u0304s(k+N\u22121) to at least one vertex of the set X\u0304s(k+N). 3.1.2 INTER-SAMPLE TRACKING In addition to the recursive feasibility requirement, to guarantee the satisfaction of the fine stage stroke constraint, y\u0304 f = re f \u2212 ys needs to be within the stroke range of the fine stage. Assuming that the terminal constraint can be satisfied (for steps k\u22121+N and k+N), as obtained from Section 3.1.1, the last piece of the coarse stage trajectory is considered to be connecting y\u0304s(k+N\u22121) to y\u0304s(k+N). This is shown in Fig. 5. It should be noted that the coarse stage is assumed to be moving linearly between two consecutive sample points. This is reasonable when the friction in (3) is small, and/or the axes experience very similar motions. The error introduced by this approximation can be computed and used to appropriately tighten the constraints. Given the motion of the coarse stage and the fine stage stroke limit (dmax, f ), the infeasible areas for tracking can be found. As it can be seen from Fig. 5, due to the significant difference between the sampling periods of the two stage (Tf << Ts) and the fact that the target reference trajectory is generated using the fine stage maximal acceleration and velocity (which are significantly higher than the one\u2019s of coarse stage), the stroke constraint can be violated even if the target reference trajectory is within the stroke range at time steps k+N\u22121 and k+N. Therefore, to guarantee stroke constraint satisfaction, the inter-sample locations of the target reference trajectory have to be examined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000165_ijaac.2018.092850-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000165_ijaac.2018.092850-Figure1-1.png", + "caption": "Figure 1 Mobile robot system with two trailers (see online version for colours)", + "texts": [ + " Conventional optimal control method cannot directly be utilised to deal with obstacle avoidance problems. Therefore, optimal control theory was proposed to integrated into vehicle path planning using an objective cost function (Wang et al., 2009; Raja and Pugazhenthi et al., 2012; Mashadi and Majidi, 2014; Xue and Sheu, 1988; Li et al., 2015). The proposed method was able to provide solutions to path planning problems with satisfied the initial/final constrains. A tractor trailer vehicle is regarded as one car-like front-steering tractor towing trailers as shown in Figure 1. With respect to the tractor, we are dealing with it like single-track two-degree-of-freedom model in order to generate the optimal free collision path according to the starting point and goal point. On the other hand, the trailers navigation will be as following, the tail of each trailer will follow the head as a target considering the vehicle head as a goal. Also, the head of each trailers will be the tail of the previous one, Section 4.2 will illustrate the navigation function for the two trailers in details" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002272_2016-01-1958-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002272_2016-01-1958-Figure7-1.png", + "caption": "Figure 7. Distortion contour at concavity 0 \u03bcm", + "texts": [ + " To consider the material property effect, a distortion analysis was performed with aluminum and a cast iron knuckles. As shown in Figure 6(a), fixed boundary conditions were applied to the bolt holes of the knuckle. The three degrees-of-freedom were fixed. To improve the reliability of the analysis, contact boundary conditions for the outer ring shield and knuckle were applied. Bolt-induced clamping torque, 180 Nm, was applied to the all three bolts shown in Figure 6(b)). A distortion analysis was performed by varying the concavity of the outer ring flange from 0 \u03bcm to 50 \u03bcm in 10 \u03bcm increments. Figure 7 shows the distortion contour at 0 \u03bcm. Maximum distortion occurs at the shield perimeter due to the size being both thin and large. It is important within this figure to focus on the outer ring distortion as key areas, that is, mounting location and raceways. Four points, outboard seal fitting point, P1, outboard raceway ball contact point, P2, inboard raceway ball contact point, P3, and inboard seal fitting point, P4, were selected from the outer ring because raceway contact points and seal mounting points are important for bearing performance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure12-1.png", + "caption": "Figure 12. Omnidirectional locomotion method.", + "texts": [ + " \u2212= =\u2032 = \u2032= \u2212 0 2/ , 1 , 1 1 l Li ui iui i ui iui ui ui i iui P P P P P PTP rr rr \u2212= =\u2032 = \u2032= \u2212 0 2/ , 1 , 1 1 r i ri iri i ri iri ri ri i iri l L P P P P P PTP rr rr ( ){ } 02/1 =\u22c5+\u2212 \u2212 uiiiui PPPP rRri +=P imi PP = { } joiwimi ttVer PMP )(, \u2212= Once the joint has moved by the wavelength of the propagation line, it stops at the position given by (11) The availability of the omnidirectional locomotion method was evaluated by the simulation of the model of the robot. In this simulation, we defined the track for the motion of the center of the model on the sphere as the desired path, and we regarded the coordinate that is determined by averaging each position of joint as the center of the model. The model was directed to move from 0 (the direction of x-axis) to 45\u00b0, as shown in Fig. 12. A wave propagation line with a 45 mm wavelength was propagated five times. The specifications for the simulation are listed in Table III. separate from the sphere during the simulation. When the robot is moving, both of the rollers of the unit cannot mechanistically adsorb onto spherical surfaces. Because we assume roller B constantly adsorb onto surfaces, roller A must separate. Therefore, we confirmed the gap between roller A and the surface. The results of the omnidirectional movement simulation are plotted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002617_10402004.2016.1271928-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002617_10402004.2016.1271928-Figure3-1.png", + "caption": "Figure 3. Schematic drawings of the setup configurations: (a) side face configuration and (b) front face configuration: 1, rotating inner ring; 2, stationary outer ring; 3, grease inside the flat-parallel gap; 4, shaft; 5, housing; 6, planes of measurement.", + "texts": [ + " The images were divided in interrogation areas with sizes as follows for the side face configuration (A): starting from 64 \u00a3 128 pixels and ending at 16 \u00a3 32 pixels; for the front face configuration (B): starting from 128 \u00a3 64 pixels and ending at 32 \u00a3 16 pixels. For more details about the mPIV technique the reader is referred to, for example, Meinhart, et al. (11) and Green, et al. (7). For all experiments the gap height was set to h D 1 mm. To assure good visualization, the rotating and stationary rings were made of polymethyl methacrylate with a roughness of Ra D 2 mm. In order to obtain images in the whole gap, for the side face configuration (Fig. 3a), the microscope focal magnification was set to 10\u00a3. On the other hand, for the front face configuration (Fig. 3b), the focal magnification was set to the minimum (5\u00a3) in order to study the largest possible radial height of the seal. mPIV test rig All experiments were carried out in the mPIV test rig (Fig. 4). It consisted of a double-pulsed Nd:YLF laser from Dantec Dynamics A/S (Copenhagen, Denmark) as a light source with a Figure 1. Conveyor idler construction. Respective parts are as follows: 1, conveyor belt; 2, idler tube; 3, bearing chamber welded to the shell; 4, rolling bearing; 5, rotating element; 6, stationary housing/element; 7, lubricating grease; 8, dust cap for additional protection; 9, idler axle", + " Grease velocity profiles in the side face (A) configuration In order to improve the optical transparency of the polymethyl methacrylate seal, the cylindrical stationary element was milled Figure 5. LG4% (side) grease velocity profiles at different depths. The rotational velocity N was kept constant at 15 rpm and the gap height h was set to 1 mm. The arrow points in the direction of increasing distance from the stationary housing to the focal plane velocity data for respective velocity profile is obtained (see Fig. 3a, item 6). from one side in order to get a flat and well-polished surface. The internal side that is in contact with the grease is also cylindrical but its diameter is very large compared to the dimensions of the actual field of view. With respect to this, it was assumed that the optical distortion is negligible. It can be added that the plots were drawn from the vectors at the center of the acquired images; that is, in the tangent direction to the cylinder surface. The grease velocity profiles in the side face configuration are shown in Figs. 5\u20138. The vertical axis represents the linear velocity of the grease and the horizontal axis represents the gap height between the rotating ring (left side) and the stationary ring (right side). The gap h was set to 1 mm and the rotational velocity N of the shaft was constant and set to 15 rpm. The nonlinear grease velocity distribution is measured at different planes starting from Fr0 (Fig. 3a, item 6) toward the rotating ring in the radial direction. Note that the Fr0 plane is situated in the interface between the grease and the stationary ring wall; in other words, at the maximal radial position (Rmax). The plots show a clear correlation between the distance from the origin (stationary housing) to the plane of Figure 7. Same as in Fig. 5 but for the LG7% grease. Figure 8. Same as in Fig. 5 but for the WSA grease. measurement and the nonlinear non-Newtonian grease velocity profile. Interestingly, the nonlinearity is found to increase with distance from the housing", + " 10\u201313, the corresponding velocity profiles in the front face configuration are shown. The vertical axis represents the radial position of the grease (0 being the contact area between the stationary outer ring and the grease mass, Rmax) and the horizontal axis represents the grease velocity. As in the previous configuration, the gap h was set to 1 mm and the velocity of the shaft N was kept constant at 15 rpm. As in the previous configuration, the grease velocity profiles were determined at different planes starting from Fx0 (Fig. 3b) toward the flat face of the Figure 11. Same as Fig. 10 but for the LG5% grease. rotating inner ring. From the plots it can be seen that the closer the determined plane of measurement is to the rotating ring, the higher the nonlinearity in the velocity profiles is. This property of the flow is in line with the observed behavior for the side faced configuration, where the shear rate increases in the direction of the rotating shaft, inducing a larger shear in the flow, which, in turn, due to the shear-thinning property of the grease, yields an increasing nonlinearity in the velocity profiles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002126_icocs.2015.7483296-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002126_icocs.2015.7483296-Figure1-1.png", + "caption": "Figure 1. Quadrotor configuration, frame system with a body fixed frame B and the inertial frame E.", + "texts": [ + "eywords: Genetic algorithm, PID, fitness, model, control, uncertain nonlinear system. I. INTRODUCTION The quadcopter is a nonlinear unstable system, which is a part of aerial robots. It is constituted by four rotors placed at the end of a cross. The center of this cross is occupied by the control circuit. Its motions are governed by six degrees of freedom: three rotations around 3 axes (x, y, z) (roll, pitch and yaw) and the three spatial translations (Figure 1)[1] [2]. The control of such system is complex, because of nonlinearity of its dynamic representation and the number of parameters, which it involves. Numerous studies have been developed to model and stabilize such systems [3] [4] [5]. The classical correction methods PID and LQ are widely used [6]. If the latter represent the advantage to be simple because they are linear, they reveal the drawback to require the presence of a linear model to synthesize. It also implies the complexity of the established laws of command because the latter must be widened on all the domain of flight of these quadcopter", + " Expressions and equations that follow are determined under some assumptions [11]: The structure of quadcopter is assumed rigid and symmetrical, allowing to consider a diagonal inertia matrix. The propellers are assumed rigid to be able to neglect deformation effect during rotation. The forces of thrust and drag are proportional to the square of the propellers speed. To describe the dynamic model of the quadcopter, we use two frames: the first is fixed and linked to the ground, and the second is mobile and linked to the studied structure (Figure 1). We find in Table 1 the used symbols and their meanings and in Table 2 the quadcopter parameters values. Note that, except the parameters m, l, b, d, Jr, I, g\u2026, given in Table 2, that are constant, the rest of the variables are timevarying. TABLE 1.SYMBOLS OF THE MODEL AND THEIR MEANING Symbol Meaning Symbol Meaning M The mass of the quadcopter \u03a9 i angular velocity of each rotor Ix, Iy, Iz The inertias around x, y and z axis Kfax, Kfay, Kfaz Frictions Aerodynamics coefficients L The half size of quadcopter Kftx, Kfty, Kftz Translation drags coefficients b, d thrust and drag coefficients , , Rotation around the roll, pitch and yaw axes Jr Rotor Inertia Fi Thrust forces of each rotor I Inertia matrix (3 \u00d7 3) of the quadcopter ,\u03a9 Linear, angular speed of the quadcopter G gravity constant (\u03a9) Skew-synmmetric matrix 978-1-4673-9669-1/15/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002290_speedam.2016.7525820-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002290_speedam.2016.7525820-Figure3-1.png", + "caption": "Fig. 3. Open angle of the rotor pole.", + "texts": [ + " A consequent pole type motor has a characteristic problem, wherein it is likely that the magnetic flux distribution becomes unstable because the magnetic characteristic of rotor magnetic poles differs from that of rotor salient poles [7]. The structure of the rotor core is also different from that of the basic model. Therefore, a relative position relationship between the stator teeth and the rotor varies. This affects the torque ripple. [4] The torque ripple may, therefore, be increased over the basic model. Fig. 3 compares the open angle of the rotor magnetic pole with Nd-Fe-B magnets a with that of the rotor salient pole a\u2019. The total volume of the auxiliary ferrite magnet and the ratio of the open angle a/a\u2019 are varied. The ratio of open angle a/a\u2019 is set at 1.05, 1.11, 1.18, and 1.25. Fig.4 shows the rotor surface section of model A when the ratio of open angle a/a\u2019 is equal to 1.18. Table IV shows the total volume of the auxiliary ferrite magnet and the ratio of open angle a/a\u2019. This creates a positional deviation of the poles between the upper core and the lower core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002815_1045389x19898251-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002815_1045389x19898251-Figure3-1.png", + "caption": "Figure 3. Working principle: (a) linear mode, (b) \u2018\u2018Arc-turn\u2019\u2019 mode and (c) \u2018\u2018U-turn\u2019\u2019 mode.", + "texts": [ + " Figure 2(a) shows the calculated vibration mode based on piezoelectric coupling, in which the calculated resonant frequency of this first-order bending mode is 200 Hz, and the maximum total displacement of the leg tip is about 0.8 mm. The displacements along X-axis and Z-axis in one cycle are plotted in Figure 2(b), in which the motion trajectory is approximate to an oblique line. This robot moves forward when legs push the ground, in which the linear motion process in one cycle can be divided into four steps, as shown in Figure 3(a): Step 1: At the excitement moment, the leg begins to push the ground from the balanced position. Point A is used as a reference point to calibrate the motion distance of the robot. Step 2: The leg lifts the robot to the maximum inclined angle a (. 75 ) at 1/4 cycle. Point A moves d1 along +X direction because of the friction force at Point B. Step 3: The leg vibrates along +X direction, reaching its balanced position at 1/2 cycle. At this moment, Point A may go back a little displacement but still be a d2 away from the original position", + " Step 4: The leg continues to vibrate along +X direction to the minimum inclined angle a (\\ 75 ), during which the reason of leaving the ground is that the vibration acceleration is much larger than gravity Then it repeats the above process for next cycle. This robot also has two types of turning modes: \u2018\u2018Arc-turn\u2019\u2019 and \u2018\u2018U-turn.\u2019\u2019 \u2018\u2018Arc-turn\u2019\u2019 mode can be driven by random three legs in L1\u2013L4, so the remaining one creates a frictional resistance which obstructs the forward movement. For example, when L1\u2013L3 are excited, but L4 is not, the robot would turn left in XOY plane, as shown in Figure 3(b) (arrow \u2018\u2018!\u2019\u2019 means the driving direction). Similarly, if L1, L2, and L4 are excited, but L3 is not, it would turn right. \u2018\u2018U-turn\u2019\u2019 mode can be driven by random one leg in L1\u2013L4. For example, when L3 is only excited in Figure 3(c), it would rotate around Point C1 near L2. Similarly, if L4 is only excited, it would rotate around Point C2 near L1. This robot can perform multi-functional locomotion through large body deformation. The contraction from the heated SMA spring realizes the lifting and falling motions of different driving legs. In Figure 4, the three rows show the heating modes and state changes of L1 and L2, L5 and L6, and L3 and L4, respectively, whereas the red part represents the heated part. Based on the structural design, the gravity center is ahead of L5 and L6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001175_tmag.2014.2356593-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001175_tmag.2014.2356593-Figure2-1.png", + "caption": "Fig. 2. Finite element model of helicopter live-line work on 1000 kV UHV transmission lines. (a) Global view. (b) Helicopter live-line work platform.", + "texts": [ + " The main parameters of line worker are displayed in Table I, which are mainly based on the statistical data recorded in GB10000-88 human dimensions of Chinese adults. At last, the modified data of line worker is used in consideration of the conductive clothing line worker wears. In consideration of the structure complexity of helicopter MD500 and the computer memory size, some simplifications of the helicopter are adopted, and the interconnecting pieces of different parts of the helicopter are ignored. The consistency of potential of helicopter is guaranteed by potential coupling, as shown in Fig. 2. The 1000 kV transmission lines in this paper adopt an eight bundle structure, the diameter of bundle conductor is 30 mm, and the bundle spacing is 400 mm. The transmission lines are in upper triangular arrangement, and due to the long line span between adjacent towers, the transmission lines are built as horizontal cylinder and in the minimum height. 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index", + " The parameters R and L are the resistance and inductance of bonding wand and conductive clothing used in the experiment of helicopter live-line work, respectively, which are deduced from measurement. Under the situation of switch K is open, the potential of live-line work platform is determined by C1 and C2, which means it has not been equipotent connected with the transmission lines. Once switch K is closed, the current through the resistance R is the discharge current. According to the finite element model built in Fig. 2 and energy method, the capacitance between helicopter live-line work platform and phase A is 118 pF, the capacitance between helicopter live-line work platform and ground is 286 pF. When the capacitances are calculated, the helicopter live-line work platform is taken as a floating electrode. The resistance R and inductance L measured by experiment are 120 and 0.1 mH. Then, the discharging current during bonding procedure can be obtained according to the equivalent circuit model, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001104_icra.2014.6907171-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001104_icra.2014.6907171-Figure2-1.png", + "caption": "Fig. 2: Contact events classes for the robotic gripper and roughneck. Labels 1 and 2 represent sensor placement locations on the robotic gripper setup detailed in Section IV-B", + "texts": [ + " Both the gripper and roughneck grasp steel pipes with hardened steel dies. Contact events (including impacts, slips, and other unspecified sources of noise) may occur at either site. The robotic gripper\u2019s primary function is to transfer pipes and move them into or out of the roughneck. During pipe transfer and placement, slips may occur along the pipe\u2019s axis. However, rotational slips about the pipe\u2019s long axis are unlikely. Contact event classes for the gripper are: (I) linear slip, (V) impact, and (VI) noise (Fig. 2). The roughneck has two sets of jaws that grip pipes in a manner similar to a lathe chuck. Its primary function is to connect sections of pipe by screwing or threading them together. Pipe threading has two stages: spin-in and makeup. During spin-in, the rotating upper jaw grips the upper pipe and spins at 100 RPM, applying torques ranging from 2- 3 kNm, while the lower jaws hold the bottom pipe stationary. If spin-in is successful, the pipe mating process culminates with a make-up operation, slowly rotating the upper pipe a bit further with a torque of 20-60 kNm", + " AE magnitudes are similar to those obtained on gripper as noted in the next section. 1) Apparatus: RDS\u2019s apparatus has a hydraulic linear actuator and a robotic gripper capable of 25.8 kN of grip force with three sets of steel dies. Experiments were performed using a 253 mm pipe. The same AE sensors were attached to various locations on the gripper, with couplant grease at the interface (Fig. 6). Grip force was held constant throughout each trial via closed-loop control. 2) Procedure: The same gripper contact events (I) linear slip, (V) impact, and (VI) noise (Fig. 2) were tested. Linear slip and impact trial protocols were unchanged with respect to the benchtop apparatus. Noise was introduced via hydraulic servo valves mounted on the system\u2019s frame. Two locations were tested (Fig. 6): location 1 is analogous to the benchtop setup in terms of AE coupling, location 2 is RDS\u2019s desired sensor placement location, as location 1 is within a hazardous \u201cred zone\u201d. At each location, three trials were conducted for impact, noise, and linear slip at two force levels (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure9-1.png", + "caption": "Figure 9. A cross-sectional diagram and side view of the fabricated initial gear prototype with flat passive rollers", + "texts": [ + " This has the advantage of flat passive rollers capable of sliding between the structure of the omnidirectional driving gear with intermittent teeth, even with only one layer of passive rollers, because the flat passive roller has more width than the conical roller in the direction of the sliding motion of the whole gear mechanism. By expanding the width of the flat structure, a flat passive roller can achieve continuous contact with the teeth of the omnidirectional driving gear while sliding. Accordingly, in the direction of the central rotating shaft, which is the same as that of the sliding motion of the whole gear mechanism, it offers the additional advantage of structural simplification. A drawing of the initial gear prototype mechanism with flat passive rollers is shown in Fig. 9, while a cross-sectional diagram of one unit of the flat passive roller and an exploded view of its components are shown in Fig. 10, and front and side views of the fabricated initial gear prototype with flat passive rollers are also shown in Fig. 11. The outline of the flat passive roller is the same as the involute curve, just like the gear with conical passive roller shown in Fig. 7. The mathematical equations of this involute curve are also equations (1) and (2), while A is 20 this time. The specifications of this initial gear prototype with flat passive rollers are shown in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003961_j.measurement.2020.108909-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003961_j.measurement.2020.108909-Figure2-1.png", + "caption": "Fig. 2. Physical object of IMD DBB.", + "texts": [ + " Hence, a Double Ballbar with enhanced measuring range was designed and manufactured at Institute of Mechatronic Engineering, TU Dresden (IMD). Unlike the common commercial Double Ballbars with transducer, IMD DBB uses an optical measuring principle. As Fig. 1 shows, the optical sensor head scans the scale tape and builds an incremental measuring system with a resolution of 0.1 \u03bcm. We experimentally validated in our work [23] that the IMD DBB could reach a comparable accuracy (1 \u03bcm in 5 mm measuring range) to the commercial DBB (\u2264 1 \u03bcm in \u00b10.1 mm measuring range by QC20W [10]). Fig. 2 visualizes the physical object of IMD DBB, the material of which is under 1000 euro cost range. Before each measurement, the DBB would be calibrated with an external calibrator (Fig. 3) and the absolute bar length refers to the calibration value. We chose Felix (Fig. 4) as the target platform for the experimental validation: a Stewart\u2013Gough platform4 with a simple 6\u20136 structure, 3 We will occasionally use the term \u2018spatial\u2019 to denote 6-D measurement, which includes not only the end effector/spindle\u2019s position in 3-D space but also the orientation thereof" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.5-1.png", + "caption": "Figure 9.5 The wrist flip phenomenon of the third kind of multiplicity.", + "texts": [ + "60) After inserting the primed angles with their expressions in terms of the unprimed angles, Eq. (9.60) can be manipulated as shown below by using the shifting property of the rotation matrices. C\u0302\u2217\u2217 = eu\u03033(\ud835\udf034+\ud835\udf0b)e\u2212u\u03032\ud835\udf035 eu\u03033(\ud835\udf036+\ud835\udf0b) = eu\u03033\ud835\udf034 eu\u03033\ud835\udf0be\u2212u\u03032\ud835\udf035 eu\u03033\ud835\udf0beu\u03033\ud835\udf036 \u21d2 C\u0302\u2217\u2217 = eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 eu\u03033(2\ud835\udf0b)eu\u03033\ud835\udf036 = eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 eu\u03033(\ud835\udf036+2\ud835\udf0b) \u21d2 C\u0302\u2217\u2217 = eu\u03033\ud835\udf034 eu\u03032\ud835\udf035 eu\u03033\ud835\udf036 = C\u0302\u2217 (9.61) As verified above, C\u0302\u2217\u2217 turns out to be the same as C\u0302\u2217. The wrist flip phenomenon can also be verified figuratively as illustrated in Figure 9.5 by means of three successive rotations. Thus, it is concluded that \ud835\udf0e5 does not lead to visually distinct poses. Therefore, it can be taken as \ud835\udf0e5 = + 1 without loss of generality. A Puma manipulator may have three distinct kinds of position singularities, which are described and discussed below. (a) First Kind of Position Singularity Equation (9.30) implies that the first kind of position singularity occurs if the position of the end-effector is specified so that r1 = r2 = 0. The same equation also implies that such a specification can be made only if the manipulator happens to be so special that d2 = 0", + " These poses are very similar to those illustrated in Figure 9.3 for the Puma manipulator. Therefore, they are also designated as left shouldered if \ud835\udf0e1 = + 1 and right shouldered if \ud835\udf0e1 = \u2212 1. They are illustrated in Figure 9.14. (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the wrist joint variables (\ud835\udf035, \ud835\udf034, \ud835\udf036). Therefore, it is the same as the third kind of multiplicity of the Puma manipulator. In other words, as illustrated in Figure 9.5, it occurs similarly as a wrist flip phenomenon without any visual distinction. A Stanford manipulator with d2 > 0 may have only one kind of position singularity. Equation (9.149) implies that it is the same as the third kind of position singularity of the Puma manipulator, which is explained in Section 9.1.5 and illustrated in Figure 9.8. (a) Angular Velocity of the End-Effector with Respect to the Base Frame Let \ud835\udf14 = \ud835\udf146. Then, Eq. (9.124) leads to the following expression. \ud835\udf14 = ?\u0307?1u3 + ?\u0307?2eu\u03033\ud835\udf031 u2 + ", + "298), because it indicates that, for the same value of c\ud835\udf032 = \ud835\udf092, \ud835\udf032 becomes positive if \ud835\udf0e2 = + 1 and \ud835\udf032 becomes negative if \ud835\udf0e2 = \u2212 1. The corresponding values of \ud835\udf031 are determined by Eqs. (9.301) and (9.302). (b) Second Kind of Multiplicity The second kind of multiplicity is associated with the sign variable \ud835\udf0e5 that arises in the process of finding the wrist joint variables (\ud835\udf035, \ud835\udf034, \ud835\udf036). Therefore, it is the same as the third kind of multiplicity of the Puma manipulator. In other words, as illustrated in Figure 9.5, it occurs similarly as a wrist flip phenomenon without any visual distinction. A Scara manipulator may have two distinct kinds of position singularities, which are described and discussed below. (a) First Kind of Position Singularity Equation (9.301) implies that the first kind of position singularity occurs if D1 = 0, i.e. if the determinant of the coefficient matrix vanishes. According to Eq. (9.300), D1 can vanish only if the following equations are both satisfied. b1 = b2 (9.311) c\ud835\udf032 = \u22121 or \ud835\udf032 = \ud835\udf0b (9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000089_cbs46900.2019.9114439-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000089_cbs46900.2019.9114439-Figure2-1.png", + "caption": "Fig. 2. The relationship between inputs (\u03b8thigh, \u03b8\u0307thigh, and \u03b8\u0308thigh on the three co-ordinate axes) and outputs (A: ankle angle, B: ankle moment, C: knee angle, and D: knee moment, whose values are represented in heat map) from the training data set. The random forest model attempts to learn this relation and then predicts on validation data set. The percentage gait cycle is also shown.", + "texts": [ + " The predicted and expected output values for each gait cycle was resampled to 100 time samples (one sample each representing one percent of gait cycle). Further, the mean RMSE of prediction was calculated for each gait cycle percentage (each of the 100 samples) across all the validation trials from all the crossvalidation iterations. Fig. 1 shows the time series measurements of the three inputs to our random forest model. This is the data for one gait cycle (time between consecutive heel contacts of right leg) as the subject performs level ground walking at self selected normal speed (mean speed = 1.28 m/s). Fig. 2 shows the input-output relationships the random forest model tries to learn from the training dataset for case II. It could be seen that there is a smooth consistent transition of the output values with respect to the three-dimensional input values across different trials, making it easier for a random forest regression model to define the input-output relationships. Table I shows the R2 values and RMSE of the random forest predictions of the ankle angle, ankle moment, knee angle, and knee moment with both the input combinations", + " This suggests that the relation between thigh angular acceleration, \u03b8\u0308thigh, and knee angle/moment provide useful information in differentiating knee angle and moment values during the start and end of the gait cycle. The reason for this could be explained as follows. In a two-dimensional input space (where inputs are \u03b8thigh, \u03b8\u0307thigh, case I), the input values overlap during the start and end of the gait cycle. This can be visualized in the thigh angle - thigh angular velocity plane (top view) of the three-dimensional plots in fig. 2. As a result, the regression model is expected to produce different output values for same range of input values, which becomes infeasible. In this scenario, adding a third dimension (\u03b8\u0308thigh, case II) to the input space can help the regressor differentiate such output values, which were not distinguishable in two-dimensional input space. For ankle angle and moment prediction, a two-dimensional input space seems to be sufficient for discriminating the output values for the whole gait cycle. Since the proposed model can predict the angles and moments required at both the ankle and the knee, it could be used to devise a control strategy for an active kneeankle prosthesis/orthosis for transfemoral amputees or stroke survivors whose residual thigh motion on the ipsilateral side can provide inputs to generate continuous control commands for actuating an active prosthesis/orthosis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure4.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure4.1-1.png", + "caption": "Figure 4.1 An aircraft observed in an earth fixed reference frame.", + "texts": [ + "13) states that the components of v\u20d7a = Dar\u20d7 in the frame a are directly equal to the derivatives of the components of the vector r\u20d7 in the same frame. On the other hand, Inequality (4.14) states that the components of v\u20d7a = Dar\u20d7 in a frameb (other than a) are not equal to the derivatives of the components of a vector in b. In other words, as stated by Eq. (4.15), the components of v\u20d7a in b are not derivatives per se but they are obtained as linear combinations of certain derivatives, which are the components of v\u20d7a in a. Consider Figure 4.1 showing an aircraft flying with respect to an earth fixed reference frame e(O). As usual, e(O) is taken as an NED (north-east-down) reference frame. Let b(C) be the body fixed reference frame attached to the aircraft, where C is the 66 Kinematics of General Spatial Mechanical Systems mass center of the aircraft. The location and orientation of the aircraft, i.e. b(C), with respect to e(O) are described by the following equations. r\u20d7C = u\u20d7(e) 1 x + u\u20d7(e) 2 y + u\u20d7(e) 3 z (4.16) r(e)C = u1x + u2y + u3z (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000986_gt2014-26756-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000986_gt2014-26756-Figure4-1.png", + "caption": "FIGURE 4. POSITIONS OF THE FILM THICKNESS SENSORS", + "texts": [ + " The dimensions of the test chamber were varied by changing the outer ring and thus the outer diameter of the chamber and/or by inserting a thicker viewing window, which would decrease the axial length of the chamber. Figure 3 depicts the four investigated bearing chamber geometries. All parameters that define an operating point, as well as several additional control parameters were permanently measured and recorded. This includes shaft speed, oil supply flow rate, air pressure, temperatures, scavenge flow rate, and air and oil outlet flows. Furthermore, the outer ring of the test chamber was equipped with capacitive film thickness sensors at nine different circumferential positions (Figure 4). A detailed description of the film thickness measurement technique is given by Kurz et al. [11]. Table 2 summarizes the uncertainties of the most important parameters. (a) b = 66 mm, h = 47 mm (b) b = 66 mm, h = 23.5 mm (c) b = 42 mm, h = 47 mm (d) b = 42 mm, h = 23.5 mm FIGURE 3. CROSS SECTIONS OF THE FOUR INVESTIGATED The regime change was identified by observing the film thickness distribution along the circumference of the bearing chamber. When the regime changes to the second regime, oil suddenly flows against the direction of gravity from the left part of the chamber (as seen in Figure 4) over the top into the right part. As a consequence the oil film thickens very abruptly in the right part of the chamber. This phenomenon was used as criterion for an objective determination of the change from regime 1 to regime 2. Since the shaft speed was the parameter which could be 3 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/80924/ on 05/01/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use changed over the widest range, all other parameters were fixed and the shaft speed was varied in steps of several hundred revolutions per minute (rpm) until the regime had changed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000005_camsap45676.2019.9022448-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000005_camsap45676.2019.9022448-Figure4-1.png", + "caption": "FIGURE 4. An illustration of airways, intersections, and nodes.", + "texts": [ + " Drones are only allowed inside the following three: airways playing a similar role to the roads, intersections formed by at least two airways, and nodeswhich are the points of interest reachable through an alternating sequence of airways and intersections. Each of these three has concrete geometric shape and is guaranteed to be collision free from static structures. Movement of drones inside the airways and intersections is regulated (for example drones must move only in the designated direction(s) of an airway or intersection) whereas inside the nodes, drones are in the free flight mode (Fig. 4). The airspace is partitioned into zones and hence each zone contains its airways, intersections, and nodes. Adjacent zones are reachable from each other through inbound and outbound gates which are the intersections at the border but they are special in that they belong to both zones. No airway is allowed to cross the border between two zones, unless it is segmented into two airways with a gate at the border joining the airways. The graph that is formed by treating both nodes and intersections (which include gates) as the vertices and airways as the directed edges is called the zone graph (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure5.6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure5.6-1.png", + "caption": "Figure 5.6 A spatial slider-crank mechanism.", + "texts": [ + "156) between the joint accelerations is of course intrinsically related to the indeterminacy expressed by Eq. (5.129) between the joint velocities. Therefore, the elimination of the indeterminacy must be consistent. For example, the elimination of the indeterminacy may be based on the continuity of \ud835\udf0301(t). Then, the values to be assigned to ?\u0307?01 and ?\u0308?01 at the instant ts of the singularity can be taken as their values immediately before the singularity at t = ts \u2212\ud835\udee5t so that ?\u0307?01(ts) = ?\u0307?01(ts \u2212 \ud835\udee5t) and ?\u0308?01(ts) = ?\u0308?01(ts \u2212 \ud835\udee5t). Figure 5.6 shows a spatial slider-crank mechanism, which is formed as a kinematic loop of four links 1, 2, 3, and 0 (the base). The joint 01 is revolute, the joint 03 is prismatic, and the joints 12 and 23 are spherical. In this example, the closure joint is selected to be 23. The following reference frames are attached to the base and the moving links. Their characterizing basis vectors are indicated in Figure 5.6. 0(O0) \u2236 O0 = O;0 = {u\u20d7(0) 1 , u\u20d7(0) 2 , u\u20d7(0) 3 } 1(O1) \u2236 O1 = A;1 = {u\u20d7(1) 1 , u\u20d7(1) 2 = u\u20d7(0) 2 , u\u20d7(1) 3 = u\u20d7(1) 1 \u00d7 u\u20d7(1) 2 } 2(O2) \u2236 O2 = B;2 = {u\u20d7(2) 1 , u\u20d7(2) 2 , u\u20d7(2) 3 } 3(O3) \u2236 O3 = D;3 = {u\u20d7(3) 1 , u\u20d7(3) 2 = u\u20d7(3) 3 \u00d7 u\u20d7(3) 1 , u\u20d7(3) 3 = u\u20d7(0) 3 } The joint frames are attached to the kinematic elements as shown below. 01(O01) = 0(A) \u2225 0(O),10(O10) = 1(A) 12(O12) = 1(B) \u2225 1(A),21(O21) = 2(B) 23(O23) = 2(C) \u2225 2(B),32(O32) = 3(C) \u2225 3(D) 03(O03) = 3(O) \u2225 3(D),30(O30) = 3(D) The nonzero constant parameters are listed below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000785_s10846-013-9973-9-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000785_s10846-013-9973-9-Figure6-1.png", + "caption": "Fig. 6 Avionics of the helicopter", + "texts": [ + ", 4 (22) where F and are the uncertainty bound matrices as | f \u2212 f\u0302 | \u2264 F (23) b = (I4\u00d74 + \u03b4)b\u0302, |\u03b4| \u2264 (24) By knowing the uncertainty bounds F and and solving Eq. 22 for Ki\u2019s at each sampling time it is guaranteed that the control outputs will follow their desired trajectories [29]. The airframe to be tested in this work is the Evolution Ex helicopter shown in Fig. 5, which is an electric model helicopter with a 2 m blade span and payload of 10 kg. The Avionics box of the helicopter is shown in Fig. 6 and includes: a PC-104 computer as the real-time embedded controller, a Crossbow NAV440 GPS-Aided Inertial Measurement Unit (IMU), an Xstream RF modem for communicating with the ground station, a Servo Switch Card (SSC) for switching between the manual radio and automatic control and a microcontroller for the battery voltage and current monitoring. The ground station computer communicates with the Avionics on the helicopter using a pair of RF modem receivers and allows for online updating and uploading the control gains as well as monitoring the attitude, rotor speed, and voltage and current of the batteries" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001506_s00170-014-6017-y-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001506_s00170-014-6017-y-Figure6-1.png", + "caption": "Fig. 6 Stress distribution of the gear tooth surface", + "texts": [ + " The maximum stress of the rack tooth surface appears on the contact surface. The stress shows symmetrical distribution along the tooth thickness and is located at the end of the contact area. However, the difference is that the maximum stress (1,034.5 MPa) on the rack tooth surface of hardening treatment is distributed in the junction between the hardening layer and the un-hardening treatment area. After hardening treatment, the stress distribution on the gear tooth surface is almost consistent before and after hardening treatment, as shown in Fig. 6. The maximum gear stress is on the tooth surface as 861.6 MPa. The tooth surface stress shows symmetrical distribution along the contact center from the contact pair and surrounding angle. 3.3 Analytical method of the lifting gear by laser processing After the laser quenching, the area perpendicular to the groove surface of the lifting gear is taken as a specimen for metallographic analysis. Through laser quenching with different process parameters, the samples are evaluated in the HXD1000TMC microhardness tester with the depth of quenching zone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003586_s12206-020-0918-5-Figure26-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003586_s12206-020-0918-5-Figure26-1.png", + "caption": "Fig. 26. Distribution of the minimum deformation.", + "texts": [ + " The deformation of the 3 - PPRU PM with an inconstant Jacobian keeps almost steady and then increases rapidly. Figs. 24 and 25, respectively, describe contour of the maximum deformation evaluation index of the PPRU 2PPRU+ PM and 3-PPRU PM. The index of the PPRU 2PPRU+ PM goes up gradually along Py-axis and keeps constant along Px-axis. Whereas the maximum deformation index of the 3-PPRU PM remain a stable trend and when the moving platform moves close to Px = \u00b10.6 m and Py = \u00b10.6 m nearby, the index rockets suddenly. Fig. 26 illustrates the minimum deformation evaluation index of the PM with a constant, partially constant and inconstant Jacobian. The PM with a completely constant Jacobian still maintains a constant deformation evaluation index about 0.8889. The output minimum deformation evaluation indices of the PM with a partially constant Jacobian and an inconstant Jacobian have a gradual decrease. Figs. 27 and 28, respectively, unfold contour of the minimum deformation evaluation index of the PPRU 2PPRU+ PM and the 3-PPRU PM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003493_j.promfg.2020.08.050-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003493_j.promfg.2020.08.050-Figure3-1.png", + "caption": "Fig. 3. Definition of the wire drawing applied in this research.", + "texts": [ + " When the internal tensile stress was increasing up to the maximum tensile strength, the wire break was occurred that affected the fine wire drawing process [1-5]. For this reason, this research aims to conduct a finite element analysis (FEA) for pure magnesium wires using the Cockroft-Latham fracture criterion to clarify the optimum drawing conditions preventing internal cracks, investigating the relationship between die half angle and pass reduction that effects on those drawability. Table 1 shows the chemical composition and the diameter of tested material wires. The schematic drawing process is shown in Fig. 3. The significate factors Do, D1, and Dn are the initial diameter of the wire, the diameter of 1st pass wire drawing, and the diameter of each passes wire drawing, respectively. These significate factors are known to affect the drawability of wire and relative to the pass reduction (R/P) and total reduction (Rt) of the wire, which are shown in equations (1) and (2), respectively [6]. The drawing conditions of a pure magnesium wire were chosen as shown in the Table 2. The conventionally available lubricant was applied on top of the material before the drawing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.20-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.20-1.png", + "caption": "Figure 9.20 First kind of position singularity of an elbow manipulator.", + "texts": [ + " That is why Link 4 appears to be folded back if \ud835\udf0e5 is switched from \ud835\udf0e5 = + 1 to \ud835\udf0e5 = \u2212 1. Based on the above explanation, it can be said that \ud835\udf0e5 = + 1 leads to an extended gripper pose and \ud835\udf0e5 = \u2212 1 leads to a folded gripper pose. An elbow manipulator may have three distinct kinds of position singularities, which are described and discussed below. (a) First Kind of Position Singularity Equation (9.205) implies that the first kind of position singularity occurs if the position of the end-effector is specified so that r1 = r2 = 0. This singularity is illustrated in Figure 9.20. As a noticeable feature of this singularity, the wrist point is located on the axis of the first joint. If this singularity occurs, \ud835\udf031 becomes indefinite and ineffective. In other words, \ud835\udf031 can be assigned an arbitrary value, but it cannot cause any change in the position of the end-effector, whatever the assigned value is. Thus, the manipulator gains a positioning freedom in the joint space, but this freedom happens to be useless in the task space as long as the wrist point is kept on the axis of the first joint", + "261) An elbow manipulator may have three distinct kinds of motion singularities, which are described and discussed below. (a) First Kind of Motion Singularity Equation (9.245) implies that the first kind of motion singularity occurs if r\u20321 = b2c\ud835\udf032 + b3c\ud835\udf0323 + b4c\ud835\udf03234 = 0 (9.262) On the other hand, referring to Part (c) of Section 9.3.2, it is seen that r\u20321 = 0 implies that r1 = r2 = 0. Therefore, the appearance of this motion singularity is the same as the appearance of the first kind of position singularity illustrated in Figure 9.20. If the manipulator has to pass through this singularity as a task requirement, then, according to Eq. (9.245), ?\u0307?1 can be kept finite by planning the task motion so that it obeys the following compatibility condition. w2c\ud835\udf031 = w1s\ud835\udf031 (9.263) If the task motion obeys the above condition, which is a singularity-induced restriction in the task space, ?\u0307?1 becomes finite but indefinite. This is because, when Eq. (9.243) is left out, which is trivially satisfied as 0 = 0, there remains five scalar equations to be satisfied by all the six joint velocities (" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001680_amm.805.105-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001680_amm.805.105-Figure6-1.png", + "caption": "Figure 6: Unit cell of an octet-truce lattice.", + "texts": [ + " The substitution of a solid volume with a less dense structure weakens the part\u2019s mechanical properties. For these experiments a structure with the highest mechanical properties is chosen. Previous experiments have shown that the octetstructure provides increased mechanical properties. Using the software netfabb by netfabb GmbH, Lupburg unit cells can be individually developed and distributed throughout a volume. For the octet structure single rods with a squared cross section and an area of 4 mm\u00b2 are used, Fig. 6. Experiments. The isolated influences of the positioning in the z-direction, the part size, the buildjob duration as well as the influence of horizontally and vertically adjacent parts on the dimensional accuracy are investigated. Thus, each specimen with one of the geometrical features is placed in a single build-job, table 2. In build-job a. the cubical specimen is placed centrally on the bottom of the building chamber and acts as the reference. In order to compare the positioning in the z-direction the specimen is positioned centrally on a maximum height of 150 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000668_fuzz-ieee.2013.6622512-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000668_fuzz-ieee.2013.6622512-Figure5-1.png", + "caption": "Fig. 5. Robot arm system", + "texts": [], + "surrounding_texts": [ + "We apply two methods in the previous section to two nonlinear systems." + ] + }, + { + "image_filename": "designv11_34_0000340_jae-141878-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000340_jae-141878-Figure1-1.png", + "caption": "Fig. 1. A magnetic gear using surface PMs. p1 = 4 pole\u2013 pairs, n2 = 17 steel poles and p3 = 13 pole-pairs on the outer rotor.", + "texts": [ + "eywords: Magnetic forces, displacement, magnetic flux density, magnetic gear Magnetic gears (MG) use a contactless mechanism for speed amplification [1\u20133]. Martin [1] proposed the coaxial MG, as shown in Fig. 1, that can achieve a high torque density. In this design there is an inner rotor, consisting of p1 pole-pair permanent magnets (PM) rotating at \u03c91, a middle rotor with n2 individual ferromagnetic steel poles that can rotate at \u03c92 and p3 pole-pair PM outer rotor rotating at \u03c93. The inner and outer rotors that contain PMs interact with the middle steel poles to create space harmonics [1\u20133]. If the relationship between the steel poles is chosen to be p1 = |p3 \u2212 n2| then the rotors will interact via a common space harmonic [1,4], the angular velocities for each rotor is then related by \u03c91 = p3 p3 \u2212 n2 \u03c93 + n2 n2 \u2212 p3 \u03c92 (1) A MG does not require gear lubrication and has the potential for high conversion efficiency [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001007_j.isatra.2014.05.028-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001007_j.isatra.2014.05.028-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of 1-DOF electrostatic torsional micromirror.", + "texts": [ + " The diameter of this micromirror is about 500 \u03bcmwith a thickness of 10 \u03bcm, and coated with 500 \u00c5/2000 \u00c5 low stress TiW/Au [7]. The mirror is suspended by two serpentine shaped hinges over a cavity etched into the Pyrex substrate. The bottom gold electrode, which usually consists of a group of individual electrodes, provides electrostatic force which its magnitude depends on the length of the electrode. The maximum tilting angle of the micromirror is designed by the mechanical stop, as the mirror touches the bottom of the cavity. A schematic diagram of the 1-DOF torsional micromirror device is shown in Fig. 2. As illustrated, a1 is the distance between the rotation axis and the nearest edge of the fixed electrode; a2 is the distance to the end of the electrode; a is the distance to the end of movable plate; b is the electrode width; and d is the vertical separation distance. The torsional motion of the micromirror can be described as follows [6]: I \u20ac\u03b8\u00feB _\u03b8\u00feK\u03b8\u00bc f \u00f0\u03b8\u00deV2 \u00f01\u00de where \u03b8 is the rotational angle, as shown in Fig. 2; I is the moment of inertia around the rotation axis and is determined by the shape of the mirror; B is the damping coefficient; K is the stiffness of the supporting hinges; V is the voltage input; and f(\u03b8) is a nonlinear function with respect to the state variable \u03b8, as [6] f \u00f0\u03b8\u00de \u00bc \u03b50b 2\u03b82 1 1 \u00f0\u03b2\u03b8=\u03b8max\u00de 1 1 \u00f0\u03b3\u03b8=\u03b8max\u00de \u00fe ln \u03b8max \u03b2\u03b8 \u03b8max \u03b3\u03b8 \u00f02\u00de which is obtained assuming small tilting angles (d\u2aa1a). Moreover, in this nonlinear function, \u03b50 is the dielectric constant of vacuum; \u03b8max \u00bc d=a is the maximum constrained tilt; \u03b2\u00bca2/a and \u03b3\u00bca1/a are constants" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002928_j.mechmachtheory.2020.103826-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002928_j.mechmachtheory.2020.103826-Figure5-1.png", + "caption": "Fig. 5. Waveboard motion.", + "texts": [ + " Note that these actions generate a force couple, denoted by M , which is congruent with the steering directions of rear and forward wheels. A physical explanation for the forward motion can be given projecting lateral forces F R and F F on the steering directions of the wheels. In Fig. 4 one can see that these forces do work, being its projections F W R and F W F nonzero, contributing to the forward propelling. Thus, rolling both decks alternately (cyclic variation of angles \u03c6r and \u03c6f ) leads to the meandering trajectory shown in sequence (1)\u2013(4) of Fig. 5 . It is important to emphasize that rear and forward forces are not applied simultaneously, existing a gap between them. In configuration (1) of Fig. 5 , the rider exerts a lateral rear force, twisting the rear platform, and the forward force is small. In order to continue with the motion, forward board is rolled in configuration (2), exerting a forefront lateral action, having the rear force become small. This sequence is repeated in configurations (3) and (4), with F R and F F being now applied in the opposite sense. The forward propelling of a wheeled vehicle as the Waveboard can be mathematically explained through the study of its predecessor system, the Snakeboard" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001534_978-3-319-01372-5-Figure3.5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001534_978-3-319-01372-5-Figure3.5-1.png", + "caption": "Fig. 3.5 Robots currently used in the CONET Integrated Testbed: five Pioneer 3-AT (left) and one Rain robot (right)", + "texts": [ + "4 Hardware 33 Three main hardware components were taken into account: platforms (static and mobile), sensors and communication infrastructure. Two main platforms are used: mobile robots and Wireless Sensor Networks, with high differences in their sensing, computing and communication capabilities. The testbed currently includes 5 skid-steer holonomic Pioneer 3-AT robots manufactured by Mobilerobots.3 The basic Pioneer 3-AT platform was enhanced with several sensors, extra computational resources\u2014a Netbook PC with an Intel Atom processor and 1024 MB SDRAM\u2014and communication equipment\u2014a IEEE 802.11 a/b/g/n Wireless bridge, see Fig. 3.5-left. The main sensors in each robot are a Microsoft Kinect, a Hokuyo laser range finder and one WSN node serially connected to the robot, which enables robot-WSN cooperation and provides extra sensing capabilities. Each robot can also be equipped with a RGB IEEE1349 camera, in case robots with two cameras are required in an experiment. Each robot is equipped with one GPS card and one Inertial Measurement Unit (IMU) for use in outdoor experiments. The testbed also includes one Rain robot developed at the Dept. Ingegneria della Informazione of University of Pisa, Fig. 3.5-right. This is a low-weight robot with differential configuration. It includes a TMoteSky node that is used as the main 3 http://www.mobilerobots.com 34 3 CONET Integrated Testbed Architecture processor of the robot. The robot includes a micro-controller responsible for lowlevel motion control. Messages to the robot are encapsulated in frames. The robot node interprets the messages and sends the corresponding commands to the microcontroller through I2C. The robot is equipped with wheel odometry sensors and five infrared distance sensors for collision detection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003737_icepds47235.2020.9249248-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003737_icepds47235.2020.9249248-Figure4-1.png", + "caption": "Fig. 4. Rotor and stator configuration of the interior permanet magnet wheelhub motor.", + "texts": [ + " The common mode or zero-sequence current has no impact to the output torque of the open-end winding machine; however, some slot effects and undesired torque pulsations are possible. The advantages of the space vector current regulation compared with sinusoidal current regulation can be summarized as follows: \u2022 15% smaller peak of the phase current flowing through semiconducting devices for the same output torque; \u2022 lower deviation of the junction temperature, which has positive impact to the aging of the switching devices; \u2022 lower maximum junction temperature for the same output torque. III. SIMULATION MODEL The sketch of the traction motor is depicted in Fig. 4. Parameters of the machine are listed in Table I. It is interior permanent magnet synchronous machine with the continuous power of 100 kW and the maximum peak power of 200 kW. The continuous torque density of the motor taking into account only active materials is 13.4 Nm/kg, which is much bigger than for the competitors [6] for continuous operating mode. Such a high torque density was granted by specific parameters of the machine, where the ohmic losses are predominant at the low speeds and the iron losses are predominant at the high speeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003347_j.matpr.2020.06.248-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003347_j.matpr.2020.06.248-Figure2-1.png", + "caption": "Fig. 2. Points and forces acting on teeth of gears.", + "texts": [ + " It specifies, for steel shafts (under definite specifications), that maximum allowable shear stress is 30% of the yield strength but not over 18% of the ultimate strength in tension for shafts without keyways as given by Aly [2] and Design of Shaft [4]. Hence, permissible shear stress, smax \u00bc 0:30Syt or 0:18Sut . Thus smax will be minimum of the two values for a safe design. Thereby, smax \u00bc 0:30 500 \u00bc 150N=mm2 or 0:18 600 \u00bc 108N=mm2. Thus, permissible shear stress is taken as smax \u00bc 108N=mm2. The points on the shaft and the forces acting on the teeth of the gears have been shown in Fig. 2 which in turn produces bending moment and torsional moment in the shaft. The forces acting on the shaft in the vertical and the horizontal planes have been shown in Fig. 3 (a) and (b) respectively along with the bending moment diagram. Using the bending moment diagrams, the values of bending moment and torsional moment at different points on the shaft have been calculated and shown in Table 2. It is now clear that the maximum value of bending moment is Mb \u00bc 18251:27N mm at point C and that of torsional moment is Mt \u00bc 6695:7N mm at point B", + " To obtain a simulation of the stresses as they occur in the component. The analysis of connecting rod was done by taking one end as a fixed support and a compressive force Fc \u00bc 4868:9N. The material taken was IS2062 Grade E250. The stresses developed is shown in Fig. 10 (a). From Fig. 10 (b), it can be seen that the connecting rod is safe to operate under the given conditions. Also, the results are similar to what was obtained through hand calculations in Section 2.2. The analysis of shaft was done by taking supports and forces as shown in Fig. 2 and Fig. 3. The material taken was AISI 4130. The results obtained are for gradual loading. The shear stresses developed is shown in Fig. 11 (a) and the factor of safety in Fig. 11 (b). It can be seen that the shaft will not fail under the given conditions and also, the results are similar to what was obtained through hand calculations in Section 2.4. Please cite this article as: A. Sinha, S. Mittal, A. Jakhmola et al., Green energy generation from road traffic using speed breakers, Materials Today: Proceedings, https://doi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001437_s0263574714001398-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001437_s0263574714001398-Figure1-1.png", + "caption": "Fig. 1. (Colour online) The 3-PRP parallel robot or ST (Star-Triangle).", + "texts": [], + "surrounding_texts": [ + "There are two main types of planar 3-PRP parallel robots. The first 3-PRP parallel robot is known as DT (double-triangle).19 The moving platform of this type is made of three beams formed like a triangle. The second planar 3-PRP parallel robot is known as ST (star-triangle) and is shown in http://journals.cambridge.org Downloaded: 16 Dec 2014 IP address: 138.251.14.35 three beams are flexible. Therefore, the moving platform of this robot is flexible. Additionally, it is assumed that the three beams have the same physical and geometrical properties. The angle between each two branches of the star is assumed to be 120\u25e6. Each of the three beams is joined to the rigid triangular base using a group of PRP joints. Finally, the rigid base is assumed to be an equilateral triangle. A general model of the ST robot in its start configuration is shown in Fig. 2(a). At the start of motion, the robot has no deformation. Center of the moving star and the center of the fixed triangular base coincide. Each of the three beams of the moving star intersects the corresponding side of the fixed base at its midpoint. To obtain an analytical model for vibration analysis, the three branches of the moving platform are considered. The three branches, also referred to as beams 1, 2, and 3, are each modeled as a discrete Euler\u2013Bernoulli beam with a prismatic joint. For beams 1, 2, and 3, a rigid body coordinate system is considered as x1w1, x2w2, and x3w3, respectively. See Figs. 2(a) and 2(b). The three rigid body coordinate systems are each attached to the rigid configuration of the corresponding beam. Center of the moving star at its rigid and deformed configurations are called G and G\u2032, respectively. Origins of the three coordinate systems are located at point G. The direction of each x-axis is along its corresponding rigid beam and passes through its revolute joint. Additionally, as shown in Fig. 2, a fixed coordinate system, XY with its origin O at the center of the equilateral triangle is defined. In this paper, in order to focus on off-load behavior of the robot\u2019s working point, point G, it is assumed that the end effector has zero concentrated inertia and experiences zero external load. As stated earlier, each branch of the moving platform is assumed to be an Euler\u2013Bernoulli beam. Therefore, the effects of the shear deformation and the rotational inertia moment are not considered in the motion equation of the beam element. Additionally, the magnitude and slope angle of the beam deformation are assumed to be small." + ] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure4-1.png", + "caption": "Figure 4. Overall view and a cross-sectional diagram of the gear with flat passive rollers", + "texts": [ + " In this gear structure, the rotational axis of each passive roller is radially aligned. The rotational axes in a radial direction have the same offset in a circular direction around the center shaft of the whole gear structure. It is easy to miniaturize and reduce the weight of this mechanism. By adapting a structure including multiple layers of passive rollers with a projection outline of an involute curve, the number of passive rollers can be reduced to at least six, as shown in Fig. 3. The second gear mechanism we examined has flat passive rollers as shown in Fig. 4. In this gear structure, the rotational axis of each passive roller is aligned continuously on the circular trajectory around the center shaft of the whole gear mechanism. Each rotational axis of a passive roller is independent. This gear mechanism eliminates the need for multiple layers of passive rollers like the gear with conical passive rollers shown in the previous chapter for sliding with an omnidirectional gear with an intermittent teeth structure. The third gear mechanism we examined has a structure resembling an ordinary commercial omnidirectional wheel with the barrel shape of passive rollers", + " In this gear structure, the passive rollers in one group between the supporting plates have a common rotational axis, while the rotational axes of passive rollers are aligned intermittently in a circular direction around the center shaft of the whole gear structure. Accordingly, this gear structure has multiple layers of passive rollers with some offset in a circular direction for continuous contact with the driven omnidirectional gear. Compared with the gear structure with flat passive rollers shown in Fig. 4, the gear structure shown in Fig. 5 can be more easily miniaturized because there are fewer rotational shafts for passive rollers. As shown above, each gear structure with passive rollers has drawbacks and advantages. Accordingly, we should choose an optimum structure according to the application of a system comprising a gear with passive rollers and an omnidirectional driving gear mechanism. The proposed gear mechanism with passive rollers has a completely different structure from previous research, whereby the direction of passive rotational axes of its small rollers was aligned with the active rotational axis of the whole gear structure [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001084_s12206-015-0217-8-Figure19-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001084_s12206-015-0217-8-Figure19-1.png", + "caption": "Fig. 19. Three-dimensional FE model of the vent line penetration nozzle.", + "texts": [ + "29 mm and 21 mm from the nozzle end, and the others were located at the inner surface region between 2.29 mm and 10.02 mm from the nozzle end [15]. Using the destructive test, it was identified that a crack penetrated through thickness of the nozzle while the others did not penetrate, as shown in Fig. 18 [15]. It was concluded that the cracks, detected in the vent line nozzle, was PWSCCs via performing the microstructure analysis using scanning electron microscope (SEM), energy dispersive X-ray spectroscopy (EDX), and transmission electron microscope (TEM) [15]. Fig. 19 depicts three-dimensional FE model of the vent line penetration nozzle. Fig. 20 shows total hoop stress distributions on the inner surface lines, calculated by applying the validated FEA procedure to the vent line penetration nozzle. As shown in Fig. 20, it is found that maximum hoop stresses at down-hill and up-hill sides occur at locations of 16.3620 mm and 17.0904 mm from point A, respectively. Location of the point A is indicated in Fig. 19. Table 4 presents maximum stress values and generation locations, and regions beyond the PWSCC initiation threshold values (210~240 MPa) [22]. Comparing the FEA results presented in Table 4 with the PWSCC initiation threshold, PWSCC initiation potentials and locations can be expected as follows: \u00b7Axial PWSCC initiation potential is much higher than circumferential PWSCC initiation potential. \u00b7Axial PWSCCs may be initiated over the inner surface region between 0 mm and 33 mm from the point A (2.9 mm to 35" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002131_oceansap.2016.7485518-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002131_oceansap.2016.7485518-Figure1-1.png", + "caption": "Fig. 1. AUV with respect to reference frame", + "texts": [ + " This paper is organized as follows: Section II highlights the modeling of AUV in the Dive Plane. In section III, the nonlinear output feedback \ud835\udc3b\u221e control algorithm is formulated, and the numerical computation is also carried out for controller design. Section IV exploits the simulation results of diving control followed by conclusion in Section V. 978-1-4673-9724-7/16/$31.00 \u00a92016 IEEE This section illustrates about the nonlinear formulation of AUV in the Dive Plane. The AUV model is a 6-degree of freedom nonlinear structure as shown in Fig. 1. The parameters of the model in the Dive Plane is considered from [13]. Thus, the state and disturbance vector are considered as \ud835\udc65(\ud835\udc61) = [\ud835\udc64, \ud835\udc5e, \ud835\udc67, \ud835\udf03]\ud835\udc47 = [\ud835\udc651, \ud835\udc652, \ud835\udc653, \ud835\udc654] \ud835\udc47 , \ud835\udc51(\ud835\udc61) = [\ud835\udc511(\ud835\udc61), \ud835\udc512(\ud835\udc61), \ud835\udc513(\ud835\udc61), \ud835\udc514(\ud835\udc61)] \ud835\udc47 , where \ud835\udc511(\ud835\udc61) and [\ud835\udc512(\ud835\udc61), \ud835\udc513(\ud835\udc61), \ud835\udc514(\ud835\udc61)] are the process noise vector and measurement noises associated with AUV dynamics. Hence, the AUV dynamics for a constant forward velocity is formulated as ?\u0307?(\ud835\udc61) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 \ud835\udc36\ud835\udc64[(\ud835\udc4a \u2212\ud835\udc35) cos\ud835\udc654 + \ud835\udc36\ud835\udc4d\ud835\udc64 \ud835\udc62\ud835\udc651 +\ud835\udc36\ud835\udc4d\ud835\udc5e \ud835\udc62\ud835\udc652 +\ud835\udc5a\ud835\udc62\ud835\udc652] \ud835\udc36\ud835\udc5e[\ud835\udc4d\ud835\udc36\ud835\udc35\ud835\udc35 sin\ud835\udc654 + \ud835\udc36\ud835\udc40\ud835\udc64 \ud835\udc62\ud835\udc651 +\ud835\udc36\ud835\udc40\ud835\udc5e \ud835\udc62\ud835\udc652] \u2212\ud835\udc62 sin\ud835\udc654 + \ud835\udc651 cos\ud835\udc654 \ud835\udc652 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc39\ud835\udc34(\ud835\udc65) + \u23a1 \u23a2\u23a2\u23a3 \ud835\udc36\ud835\udc64\ud835\udc36\ud835\udc511 0 0 0 \ud835\udc36\ud835\udc5e\ud835\udc36\ud835\udc512 \ud835\udefc 0 0 0 0 \ud835\udefc 0 0 0 0 \ud835\udefc \u23a4 \u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc3a\ud835\udc51(\ud835\udc65) \ud835\udc51(\ud835\udc61) + \u23a1 \u23a2\u23a2\u23a3 \ud835\udc36\ud835\udc64\ud835\udc36\ud835\udc4d\ud835\udeff\ud835\udc60 \ud835\udc36\ud835\udc5e\ud835\udc36\ud835\udc40\ud835\udeff\ud835\udc60 0 0 \u23a4 \u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc3a\ud835\udc62(\ud835\udc65) \ud835\udc62(\ud835\udc61) (1) \ud835\udc67(\ud835\udc61) = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 \ud835\udc4a1 0 0 0 0 \ud835\udc4a2 0 0 0 0 \ud835\udc4a3 0 0 0 0 \ud835\udc4a4 0 0 0 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc3b1(\ud835\udc65) \ud835\udc65(\ud835\udc61) + \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 0 0 0 0 \ud835\udc4a\ud835\udc62 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc3f12(\ud835\udc65) \ud835\udc62(\ud835\udc61) (2) \ud835\udc66(\ud835\udc61) = \u23a1 \u23a30 1 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc3b2(\ud835\udc65) \ud835\udc65(\ud835\udc61) + \u23a1 \u23a30 \ud835\udefc 0 0 0 0 \ud835\udefc 0 0 0 0 \ud835\udefc \u23a4 \u23a6 \ufe38 \ufe37\ufe37 \ufe38 \ud835\udc3f21(\ud835\udc65) \ud835\udc51(\ud835\udc61) (3) where \ud835\udc62(\ud835\udc61) = \ud835\udeff\ud835\udc60, \ud835\udc36\ud835\udc64 = (\ud835\udc5a\u2212 \ud835\udc36\ud835\udc4d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002296_acc.2016.7526823-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002296_acc.2016.7526823-Figure2-1.png", + "caption": "Fig. 2. Illustration of distances and bearing angles between unicycles.", + "texts": [ + " As in [3], each unicycle i establishes its local coordinate frame, i.e., the Frenet-Serret frame, with the origin at its position pi and the x-axis coincident with its orientation \u03b8i. Then, convert the states [pT j \u03b8j ] T, j \u2208 N\u0304i(\u03d5), i \u2208 O, measured in the inertial frame to those in the Frenet-Serret frame of escort i by using the following coordinate transformation [24]: xji yji \u03b8ji = cos \u03b8i sin \u03b8i 0 \u2212 sin \u03b8i cos \u03b8i 0 0 0 1 xj \u2212 xi yj \u2212 yi \u03b8j \u2212 \u03b8i . (2) As shown in Fig. 1(a) and Fig. 2, each escort i can obtain the relative positions [xji yji] T, j \u2208 N\u0304i(\u03d5), and the separation angles \u03d5(i+)i and \u03d5i(i-) by measuring the relative orientations \u03b8ji, the relative distances dji and the bearing angles \u03b2ji, j \u2208 N\u0304i(\u03d5). In fact, \u03d5(i+)i, \u03d5i(i-), dji and \u03b2ji vary with the relative positions between each escort and the target. According to [3], the evolution of the forward separation angle \u03d5(i+)i is governed by \u03d5\u0307(i+)i = \u03b7(i+)\u2212 \u03b7(i), (3) where \u03b7 : O 7\u2192 R is defined as \u03b7(j) = (vj sin\u03b20j + v0 sin\u03b2j0)/d0j ", + " Function \u03c3 : R 7\u2192 R is a C1 function such that (i) \u03c3(0) = 0, (ii) |\u03c3(z)| \u2264 1, \u2200z \u2208 R and (iii) z\u03c3(z) > 0, \u2200z \u2208 R/0, for instance, \u03c3(z) = tanh(z) and \u03c3(z) = z/ \u221a 1 + z2. Moreover, \u03c8i is defined as \u03c8i = \u03c6i \u2212 \u03b8i, (18) which denotes the direction of ~ui measured in the local coordinate frame of escort i, i.e., the vectorial angle by rotating ~ui counterclockwise until its direction coinciding with that of ~vi. Then, \u03c8i can be calculated with cos\u03c8i = (v2i + u2i \u2212 v20)/(2viui). (19) Furthermore, under assumption [A3], each escort i can use \u03b2ji and \u03b8ji to calculate \u03b2ij , see Fig. 2. Moreover, in order to obtain the separation angles \u03d5(i+)i and \u03d5i(i-) which vary with p\u0303, each escort i can use \u03b20i, \u03b2i0, \u03b2ki, \u03b2ik, d0i and dki, k \u2208 Ni(\u03d5), to calculate d0k, \u03b2k0 and \u03b20k, as shown in Fig. 2. Then, \u03d5(i+)i and \u03d5i(i-) can be obtained, and it follows from (4) that \u03b7(j), j \u2208 N\u0304i(\u03d5), can be obtained by escort i. Hence, controller (14)-(15) with (16)-(17) can be described in the form of (5). Now, the main result of this paper is stated as follows. Theorem 3.1: Consider N + 1 unicycles in the form of (1) under assumptions [A1]-[A3]. Define ui, i = 1, ..., N , as in (12). If the initial condition ui(t0) \u2265 c + k2r, \u2200t0 \u2265 0, is satisfied, controller (14)-(15) with (16)-(17) solves the escorting and patrolling control problem " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.10-1.png", + "caption": "Figure 9.10 Second kind of motion singularity of a Puma manipulator.", + "texts": [ + "81), which are to be satisfied by three joint velocities ?\u0307?1, ?\u0307?2, and ?\u0307?3. In other words, as an expected singularity feature, this motion singularity also induces a motion freedom in the joint space. Thus, an arbitrary value can be assigned to ?\u0307?1 and the others (?\u0307?2 and ?\u0307?3) can then be found as explained in Part (a) of Section 9.1.7. (b) Second Kind of Motion Singularity Equations (9.86) and (9.87) imply that the second kind of motion singularity occurs if cos \ud835\udf033 = 0 or \ud835\udf033 = \ud835\udf0e\u20323\ud835\udf0b\u22152 with \ud835\udf0e\u20323 = \u00b11 (9.99) This singularity is illustrated in Figure 9.10 in its extended version (\ud835\udf033 = \ud835\udf0b/2) and folded version (\ud835\udf033 = \u2212\ud835\udf0b/2). Note that the appearance of the folded version of this motion singularity becomes the same as the appearance of the second kind of position singularity if d4 = b2. In this singularity, with \ud835\udf0323 = \ud835\udf032 + \ud835\udf0e\u20323\ud835\udf0b\u22152, Eqs. (9.79) and (9.81) can be manipulated into the following forms. (b2?\u0307?2 + \ud835\udf0e\u20323d4?\u0307?23)s\ud835\udf032 = \u2212w\u2217 1 (9.100) (b2?\u0307?2 + \ud835\udf0e\u20323d4?\u0307?23)c\ud835\udf032 = \u2212w3 (9.101) The above equations become consistent and thus they can give finite values for " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure11-1.png", + "caption": "Figure 11. Perspective, front and side views of the fabricated initial gear prototype with flat passive rollers", + "texts": [ + " Accordingly, in the direction of the central rotating shaft, which is the same as that of the sliding motion of the whole gear mechanism, it offers the additional advantage of structural simplification. A drawing of the initial gear prototype mechanism with flat passive rollers is shown in Fig. 9, while a cross-sectional diagram of one unit of the flat passive roller and an exploded view of its components are shown in Fig. 10, and front and side views of the fabricated initial gear prototype with flat passive rollers are also shown in Fig. 11. The outline of the flat passive roller is the same as the involute curve, just like the gear with conical passive roller shown in Fig. 7. The mathematical equations of this involute curve are also equations (1) and (2), while A is 20 this time. The specifications of this initial gear prototype with flat passive rollers are shown in Table 2. In the following chapters, we will describe the results of experiments performed to confirm the features and advantages of two fabricated prototypes of gears with two different types of passive rollers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001127_humanoids.2014.7041365-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001127_humanoids.2014.7041365-Figure3-1.png", + "caption": "Fig. 3. Left: Motion of the 2 DOF planar robot along the first PMD that defines a positive coupling between the two joint angles \u03b81 and \u03b82, i.e. both joints move in the positive (counterclockwise) sense. Right: Motion along the second PMD that defines a negative coupling, i.e. \u03b81 moves counterclockwise while \u03b82 moves clockwise.", + "texts": [ + " an identity mapping matrix indicates that all the degrees of freedom are independently actuated, a mapping matrix with some zero rows indicates that some degrees of freedom are not actuated (i.e. fixed), and a mapping matrix with columns with several non-zero elements indicates that some degrees of freedom are coupled, like in the case of PMDs. Let consider a 2-DOF planar robot with two rotational joints, and let define a coupling between them using the following two PMDs: u1 = 1\u221a 2 [1, 1]T \u03bb1 = 1.0 (8) u2 = 1\u221a 2 [1,\u22121]T \u03bb2 = 0.1 (9) i.e. the first PMD defines a positive coupling between joints, as shown in Fig. 3 (left) and the second one a negative coupling, as shown in Fig. 3 (right). Fig. 4 shows the solutions found for a given query in an environment with a single obstacle using the standard RRT algorithm (top row) and the RRT\u2217 optimizing distance (bottom row) with an increasing number of samples. Observe that, as reported in [7], the RRT solution does not improve with an increasing number of samples while RRT\u2217 does. For the case of an obstacle-free environment, Fig. 5 shows a query that goes from the center of the configuration space towards the top-left corner, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001680_amm.805.105-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001680_amm.805.105-Figure8-1.png", + "caption": "Figure 8: Dimensional deviation using optical fringe pattern projection with a solid cube.", + "texts": [ + " The dimensional homogeneity across the whole specimen and in every direction is a major indicator for dimensional accuracy. Thus, the range which represents the absolute distribution of results also represents the reproducibility. The range is the difference between the maximum and the minimum value of a dataset and thus is very sensitive to single outliers. In the present chapter the two measurement methods are compared. Then the results of all influences are shown and eventually validated by the results from the multibody build-jobs. Comparison of Measurement. Fig. 8 depicts a false color image of a solid cube specimen using the fringe pattern projection method. The relative dimensional deviation is shown by the false color where red indicates an oversize and blue an undersize. On the cube\u2019s rear certain layer marks can be observed as they are process inherent. By using exemplarily six cubes with certain oversize between the bottom and top surfaces the tactile measuring method is compared to the fringe pattern projection method as shown in Fig. 9. It can be seen that the measurement methods only deviate with 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003916_s42853-020-00082-7-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003916_s42853-020-00082-7-Figure1-1.png", + "caption": "Fig. 1 Electric utility terrain vehicles. a Mechron 2230 (http:// utv.daedong.co.kr/main.do). b Hisun Sector E1 (https://www. hisunmotors.com/)", + "texts": [ + " For agricultural UTVs operated by internal combustion engines, most noises are generated by engine explosion sound; thus, to reduce the noise of agricultural UTVs, it is important to reduce engine noise. Recently, the demand for the development and supply of an eco-friendly agricultural machinery has increased. Thus, the electrification of agricultural powertrain has been discussed (Kim et al. 2014). In particular, because agricultural UTVs and simple powertrain systems require small power for work or operation, electric UTVs have been widely studied and spread to replace UTVs with internal combustion engines (Fig. 1). Noise in electric powertrain is mainly generated by the drive motor and gearbox, and it exhibits high-frequency characteristics compared to the noise of the engine powertrain (Bassett et al. 2014). Therefore, it is important to reduce the noise of the gearbox to reduce the noise of electric agricultural UTVs. The causes of noise existing in the gearbox include the following: whine noise owing to gear meshes, rattle noise owing to torque fluctuations, and noise generated by rotating shafts or bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001482_1.4029187-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001482_1.4029187-Figure6-1.png", + "caption": "Fig. 6 A dead center position for the double butterfly eight-bar linkage", + "texts": [ + " The coupler GH of the equivalent four-bar linkage may be regarded as being rigidly attached to CDEF, and A0G to A0AA0, and B0H to B0BB0. According to the criterion, in this case, the dead center position occurs when the three passive joints I23, I34, and I14 lie a common line, as shown in Fig. 5. Example 5. Dead center positions for the double butterfly eightbar linkage. Foster and Pennock [18] proposed a geometric method to locate the secondary instant centers of the double butterfly eight-bar linkage (Fig. 6), which consists of 3 five-bar loops. The double butterfly linkage in Fig. 6 has the similar dimensions as Table 1 in Ref. [18]. The dead center positions of the double butterfly linkage can be determined by any of its equivalent linkages. Let links 1, 2, 3, and 7 as the four original links from the double butterfly linkage and they are put in order as the reference link, the input link, the couple link, and the output link. Thus, the corresponding equivalent linkage I12I23I37I71 comprised the instant centers of two neighboring links. If the three passive joint I23, I37, and I71 of the equivalent linkage lie on a common straight lie, the equivalent linkage is at the dead center positions and hence the whole linkage must be at dead center positions (Fig. 6), which can be verified by Eqs. (1) and (2). Example 6. Dead center positions for the Stephenson type III six-bar linkage. A Stephenson type III six-bar linkage contains a four-bar loop and a five-bar loop, as shown in Fig. 7. With the input given through the link 2 or the joint I23, the Stephenson type III six-bar contains two types of equivalent four-bar linkages. The type I four-bar linkage consists of the four-bar loop, or the four instant centers I12, I23, I34, and I41. The type II four-bar linkage consists of the original four links as link 1, link 2, link 5, and link 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002983_b978-0-12-818546-9.00016-6-Figure11.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002983_b978-0-12-818546-9.00016-6-Figure11.2-1.png", + "caption": "Figure 11.2 Board-based canes: (A) rolling cane [59]; (B) step-up cane [60].", + "texts": [ + " On the other hand, people suffering from hemiplegia may have difficulties with quad canes and discover straight canes with higher efficiency. Quad canes help patients to stand up lonely without extra help [56,57]. Able Tripod is a type of board-based cane in which a soft triangular tip is employed to improve the connection with ground and present higher stability on several kinds of ground such as icy and slippery floors [58]. Rolling canes are designed with a three-caster base, which present the functionalities of quad canes excluding lifting canes during walking (Fig. 11.2A). This Chapter 11 Assistive devices for elderly mobility and rehabilitation: review and reflection 309 type of canes is also equipped with a braking system for stability improvement. Another type of board-based cane, pilot stepup cane, is assembled with a flip-up to enhance users\u2019 stability during stair climbing and curb (Fig. 11.2B). A smart cane was presented with navigation and healthmonitoring functions for eldercare usages [61]. A set of sonar sensors was mounted in front of an omnidirectional base for obstacle detection and localization. Au et al. [62] utilized a set of sensors in a conventional cane including accelerometer, gyroscopes, and force sensors to record the user\u2019s characteristics for decreasing the risk of fall during ambulation. According to the recorded information, the smart cane sends feedback through a speaker to guide the user to modify his/her movement for fall prevention" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001898_6.2013-5013-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001898_6.2013-5013-Figure1-1.png", + "caption": "Figure 1. Generic tail-controlled missile with body fixed \ud835\udc35-frame [9]", + "texts": [ + " To demonstrate the robustness with respect to modeling uncertainties and actuator constraints, a realistic parameter and measurement uncertainty spectrum is chosen. This paper is organized as follows. Section II describes the 6DOF simulation model of the generic tail-controlled missile. The control approach is presented in Section III. Simulation results for demanding STT-maneuvers exploiting the full physical capabilities of the missile are presented in Section IV. Finally, Section V summarizes concluding remarks. II. Simulation Model of the Missile The generic tail-controlled air defense missile considered within this paper is depicted in Figure 1. The following section presents a short description of the missile dynamics. A more detailed representation of the missile model can be found in [9]. For the purpose of evaluating the designed control algorithm, the most challenging part of the missile flight phase is considered: the terminal phase. Within this flight phase, high demands in terms of agility, performance, and tracking accuracy are posed on the missile system - especially the flight control system (FCS). The booster is assumed to be burned out by the beginning of this flight phase [14]", + " Therefore, no propulsion force acts on the missile. Since an air defense missile covers only a relatively short distance within the endgame, it is legitimate to consider the earth as non-rotating and flat. With this assumption, two main coordinate frames are sufficient to describe the missile\u2019s rigid body dynamics: a body-fixed frame (\ud835\udc35-frame) and an Earth Centered Earth Fixed frame (\ud835\udc38-frame). The body-fixed frame is attached to the missile\u2019s center of gravity (c.g.) with the x-axis pointing to the missile\u2019s cone (see Figure 1). The \ud835\udc38-frame, which is used to describe the missile\u2019s position dynamics, has its center located at sealevel with the \ud835\udc65- and \ud835\udc66-axis pointing in north- and east-directions, respectively, while the \ud835\udc67-axis points downward, perpendicular to the earth\u2019s surface. By using the feasible assumptions of a flat and non-rotating earth, the missile\u2019s velocity (\ud835\udc7d\ud835\udc3e \ud835\udc3a)\ud835\udc35 \ud835\udc38 = [\ud835\udc62\ud835\udc3e \ud835\udc3a \ud835\udc63\ud835\udc3e \ud835\udc3a \ud835\udc64\ud835\udc3e \ud835\udc3a]\ud835\udc35 \ud835\udc38,\ud835\udc47 results from the translational dynamics (?\u0307?\ud835\udc3e \ud835\udc3a) \ud835\udc35 \ud835\udc38\ud835\udc35 = \ud835\udc74\ud835\udc35\ud835\udc38 \u22c5 (\ud835\udc88 \ud835\udc3a)\ud835\udc38 + 1 \ud835\udc5a \u22c5 (\ud835\udc6d\ud835\udc34 \ud835\udc3a)\ud835\udc35 \u2212 (\ud835\udf4e\ud835\udc3e \ud835\udc38\ud835\udc35)\ud835\udc35 \u00d7 (\ud835\udc7d\ud835\udc3e \ud835\udc3a)\ud835\udc35 \ud835\udc38 (1) with (\ud835\udc88\ud835\udc3a)\ud835\udc38 and (\ud835\udc6d\ud835\udc34 \ud835\udc3a)\ud835\udc35 describing the gravitational acceleration and the aerodynamic forces, respectively", + " \u22c5 \ud835\udc46\ud835\udc5f \u22c5 [ \ud835\udc36\ud835\udc650(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) + \ud835\udc36\ud835\udc65\ud835\udf09(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e, \ud835\udf09) \ud835\udc36\ud835\udc660(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) + \ud835\udc36\ud835\udc66\ud835\udf01(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e, \ud835\udf01) \ud835\udc36\ud835\udc670(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) + \ud835\udc36\ud835\udc67\ud835\udf02(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e, \ud835\udf02) ] \ud835\udc35 (6) and moment coefficients (\ud835\udc74\ud835\udc34 \ud835\udc3a)\ud835\udc35 = ?\u0305? \u22c5 \ud835\udc46\ud835\udc5f \u22c5 \ud835\udc59\ud835\udc5f \u22c5 [ \ud835\udc36\ud835\udc3f \ud835\udc36\ud835\udc40 \ud835\udc36\ud835\udc41 ] \ud835\udc35 = ?\u0305? \u22c5 \ud835\udc46\ud835\udc5f \u22c5 \ud835\udc59\ud835\udc5f \u22c5 [ \ud835\udc36\ud835\udc3f0(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) + \ud835\udc36\ud835\udc3f\ud835\udc5d(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) \u22c5 \ud835\udc59\ud835\udc5f\ud835\udc52\ud835\udc53 2\ud835\udc49\ud835\udc3e \ud835\udc3a \u22c5 (\ud835\udc5d\ud835\udc3e \ud835\udc38\ud835\udc35)\ud835\udc35 + \ud835\udc36\ud835\udc3f\ud835\udf09(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) \u22c5 \ud835\udf09 \ud835\udc36\ud835\udc400(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) + \ud835\udc36\ud835\udc40\ud835\udc5e(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) \u22c5 \ud835\udc59\ud835\udc5f\ud835\udc52\ud835\udc53 2\ud835\udc49\ud835\udc3e \ud835\udc3a \u22c5 (\ud835\udc5e\ud835\udc3e \ud835\udc38\ud835\udc35)\ud835\udc35 + \ud835\udc36\ud835\udc40\ud835\udf02(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) \u22c5 \ud835\udf02 \ud835\udc36\ud835\udc410(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) + \ud835\udc36\ud835\udc41\ud835\udc5f(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) \u22c5 \ud835\udc59\ud835\udc5f\ud835\udc52\ud835\udc53 2\ud835\udc49\ud835\udc3e \ud835\udc3a \u22c5 (\ud835\udc5f\ud835\udc3e \ud835\udc38\ud835\udc35)\ud835\udc35 + \ud835\udc36\ud835\udc41\ud835\udf01(\ud835\udefc\ud835\udc3e , \ud835\udefd\ud835\udc3e , \ud835\udc40\ud835\udc4e) \u22c5 \ud835\udf01 ] \ud835\udc35 . (7) The air-defense missile considered within this paper is controlled via four aerodynamic fins attached to the missile\u2019s rear section (Figure 1). Each of the four fins denoted with \ud835\udeff\ud835\udc56, \ud835\udc56 = 1, \u2026 ,4 can be deflected independently and include a sensor system, which measures the angular deflection. The fins are modeled as a dynamical system of the order two, including limits for deflection \ud835\udeff\ud835\udc56,\ud835\udc59\ud835\udc56\ud835\udc5a\ud835\udc56\ud835\udc61 , deflection rate ?\u0307?\ud835\udc56,\ud835\udc59\ud835\udc56\ud835\udc5a\ud835\udc56\ud835\udc61 , and deflection acceleration ?\u0308?\ud835\udc56,\ud835\udc59\ud835\udc56\ud835\udc5a\ud835\udc56\ud835\udc61. Since the D ow nl oa de d by R O K E T SA N M IS SL E S IN C . o n Ja nu ar y 8, 2 01 5 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 3- 50 13 American Institute of Aeronautics and Astronautics 5 aerodynamic data set (6), (7) is expressed in the aerodynamic equivalent controls \ud835\udc96 = [\ud835\udf09 \ud835\udf02 \ud835\udf01]\ud835\udc47 , these entities serve as inputs for control design purposes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003852_0954406220978269-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003852_0954406220978269-Figure1-1.png", + "caption": "Figure 1. 2-DOF robotic manipulator.", + "texts": [ + " The dead zone technology is _\u0302a0 \u00bc u0ksk; if ksk > K 0; otherwise ; ( _\u0302a1 \u00bc u1kskkqk; if ksk > K 0; otherwise ; ( _\u0302a2 \u00bc u2kskkqk2; if ksk > K 0; otherwise ( (34) where K is a small positive constant. Remark 9: The selection of K is very important, it will affect the control performance of the closed-loop system. Because of the inherent robustness of the sliding mode control method, the parameter of K can not be too small. In this section, we will demonstrate the trajectory tracking performance of the proposed control strategy by the numerical simulation on 2-DOF robotic manipulator. As shown in Figure. 1, a robotic manipulator with 2 revolute joints is mounted on the base. For the dynamics model (1), we have M\u00f0q\u00de \u00bc p1 \u00fe p2 \u00fe 2p3cosq2 \u00fe J1 p2 \u00fe p3cosq2 p2 \u00fe p3cosq2 p2 \u00fe J2 C\u00f0q; _q\u00de \u00bc p3 _q2sinq2 p3\u00f0 _q1 \u00fe _q2\u00desinq2 p3 _q1sinq2 0 G\u00f0q\u00de \u00bc p4gcosq1 \u00fe p5gcos\u00f0q1 \u00fe q2\u00de p5gcos\u00f0q1 \u00fe q2\u00de where p1 \u00bc \u00f0m1 \u00fem2\u00del21; p2 \u00bc m2l 2 2; p3 \u00bc m2l1l2; p4 \u00bc \u00f0m1 \u00fem2\u00del1; p5 \u00bc m2l2. The system parameters of the robotic manipulator are listed in Table 1. The frictional effects in the joints are ignored. Considering that the uncertainties and sudden disturbances may be involved in the closedloop system, we can assume that the uncertainties are DM \u00bc 0:05M0; DC \u00bc 0:05C0; DG \u00bc 0:05G0 and the sudden disturbances are given as sd \u00bc 10sin\u00f02t\u00de \u00fe 5sin\u00f0200pt\u00de 10cos\u00f02t\u00de \u00fe 5sin\u00f0200pt\u00de ; t 3s In order to verify the superiority of the proposed method, the simulation results are compared with two other types of sliding mode controllers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000712_detc2013-12492-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000712_detc2013-12492-Figure5-1.png", + "caption": "Figure 5. EXAMPLES FROM DEVELOPMENT OF WATER BOTTLE REFILLING STATION.", + "texts": [ + " Tom\u2019s passion for changing the world caught the attention of many students at the Project Kickoff, and a capstone team of five mechanical engineers was paired with a team of five junior-level honors business students (the same course pairing for Maximus V) to undertake the project. The team surveyed numerous potential customers to understand what they would like to see in such a water bottle refilling station and then developed numerous concepts to meet these customer needs. Patent searches were conducted to assess the IP landscape both with respect to the design of the station itself as well as its internal components (e.g., water bottle cleaning). Figure 5 shows examples from prototype developing including early concept sketches, CAD models, and structural analysis. By the end of the semester, the team had developed detailed assembly drawings along with a complete bill of materials and cost estimates for construction of each water bottle refilling station. They also built a working prototype that would recognize, clean, and refill a single water bottle while playing a 30-second advertisement (see Figure 6). The efforts of the Fall 2009 capstone team were recognized with a 1st Place Best Project award as determined by industry judges who evaluated the 41 capstone projects that were at the Design Showcase that semester" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001124_icra.2013.6630772-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001124_icra.2013.6630772-Figure8-1.png", + "caption": "Figure 8. Perspective, front and side views of the fabricated initial gear prototype with conical passive rollers", + "texts": [ + " The origin of the x-y coordinate is placed on the starting point of the involute curve of the conical passive roller at its bottom. The number of teeth in this initial gear prototype is 32, hence the value of A is also 32. The combination of the precise digital drawing according to the above equations with the three-dimensional CAD system of SolidWorks 2010 and the high manufacturing function of the CNC milling machine controlled with the G-code as the output data from SolidWorks enabled the structures of gears with conical and flat passive rollers, as shown in Fig. 8 and 11. The specifications of this initial gear prototype with conical passive rollers are shown in Table 1. During this research, we also developed a gear with flat passive rollers. This has the advantage of flat passive rollers capable of sliding between the structure of the omnidirectional driving gear with intermittent teeth, even with only one layer of passive rollers, because the flat passive roller has more width than the conical roller in the direction of the sliding motion of the whole gear mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001946_ecce.2014.6953971-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001946_ecce.2014.6953971-Figure2-1.png", + "caption": "Fig. 2 Finite elements analysis", + "texts": [ + " Although the LS model gives a good description of the magnetic losses, the generalized Bertotti model could be more suitable because it is faster and easier to calibrate with the material type. The developed model for our application obtains the FDW from a nonlinear nodal network with the methodology described in [11] and solving the problem in magneto-static. The IPMSM of Fig. 1 can be discretized in different areas as described on Fig. 3. These areas represent the main flux ways as described in Fig. 2 which have been defined from a preliminary study of the IPMSM with finite elements software (FLUX 2d). The nodal network has been defined from a first study with FE approach is given in Fig. 4. So this model cannot provide the temporal FDW in the different parts of the stator but the spatial FDW. However it is possible to assume that the temporal FDW in the teeth and the stack are identical with the spatial FDW. With this assumption, only the slot harmonics are neglected. For flux density variation in the rotor, it is only caused by the harmonics linked to slots effects and inverter modulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002961_0036850419897221-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002961_0036850419897221-Figure3-1.png", + "caption": "Figure 3. Simplified fluid solid coupling model.", + "texts": [ + " The method used in this article is to establish a two-way fluid\u2013solid coupling model between brush bristle and flow field to numerically calculate the change in shape and position of brush bristle during the constant pressure difference working process, and then to calculate the brush bundle thickness and porosity. Thus, the viscous drag coefficient and inertial drag coefficient can be calculated to determine the key parameters in the porous media model. Fluid\u2013solid coupling model of brush seal Because of the large number of circumferentially symmetric bristles, a simplified model is used here. The model includes two rows and six columns of bristle, and the model size is shown in Figure 3. The material used for brush is nickel-based super alloy, its elasticity modulus is 213.7GPa, its Poisson\u2019s ratio is 0.29. Under actual working conditions, brush seal would have blown down effect that affects the effective sealing gap due to the radial runout of the rotating shaft, excessive wear, and timely change of working conditions. Only the blow down effect caused by the back baffle is considered in this article, which will be specifically analyzed below. The modeling process is as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000856_tmag.2014.2362515-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000856_tmag.2014.2362515-Figure7-1.png", + "caption": "Fig. 7. Maximum magnetic field strength distribution of a part of the stator core. (a) Non-stress. (b) Model 001. (c) Model 002. (d) Model 003.", + "texts": [ + " The local loci of Bx \u2212 By and Hx \u2212 Hy and the hysteresis loops of the x and y components with the assumed stress differed in comparison with those of the nonstress. The rotating magnetic flux in the central part of the teeth occur as shown in Fig. 5(b), and the alternating magnetic flux in the edge of the teeth and core back occur, as shown in Fig. 5(a) and (c). The change of the loci of Hx \u2212 Hy became larger than that of Bx \u2212 By . Fig. 6 shows the calculated maximum magnetic flux density distribution of each models. The maximum magnetic flux density distribution is hardly changed by a change in the assumed local stress. Fig. 7 shows the calculated maximum magnetic field strength distribution of each model. The change in the maximum magnetic field strength distribution became larger than that of the maximum magnetic flux density distribution. It was clarified that the magnetic field strength vector is more easily affected than the magnetic flux density vector due to the stress. Fig. 8 shows the calculated magnetic power loss distribution of each model. The magnetic field strength and magnetic power loss of the central part of the teeth in the stator core decrease due to the uniaxial tensile stress along the longitudinal direction of the teeth, as shown in Fig. 7(b) and (d). In addition, the magnetic field strength and magnetic power loss decreased due to the biaxial tensile stress, as shown in Fig. 7(c). Fig. 9 shows the average value of the maximum magnetic field strength Bave, the maximum magnetic field strength Have, and the magnetic power loss Wave of the stator core. The value of Bave hardly changes with and without the uniaxial and biaxial tensile stress. On the other hand, the value of Have and Wave decreased when applying the assumed local stress. In particular, the change in Have and Wave for model 001 was largest because the magnetic properties improve by applying the parallel tensile stress to the direction of the B vector in the teeth and the core back part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002290_speedam.2016.7525820-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002290_speedam.2016.7525820-Figure1-1.png", + "caption": "Fig. 1. Structure of basic model.", + "texts": [ + " Accordingly, we have discussed the development of PMSMs using Nd-Fe-B and ferrite magnets [4] [5]. In this paper, we propose novel rotor structures using rare earth and ferrite magnets. The total volume of rare earth materials used in our proposed model can be reduced by 57.5 % compared with that of the basic model. The torque characteristics of the basic and proposed models are analyzed using the three-dimensional finite element method. We demonstrate that one of our proposed models has the same torque characteristics as that of the basic model. II. BASIC MODEL Fig. 1 shows the structure of the basic model with only NdFe-B magnets. This model has eight poles and 36 slots. The total volume of the Nd-Fe-B magnets is 58000 mm3. Tables I and II show the specifications of the basic model and Nd-Fe-B magnet, respectively. 978-1-5090-2067-6/16/$31.00 \u00a92016 IEEE III. PROPOSED MODELS Fig. 2 shows a new rotor structure, and the red arrows show the magnetized direction. Model A comprises two claw-pole-type iron cores which are called the upper and lower cores, a shaft of non-magnetic material, Nd-Fe-B magnets, a cylindrical ferrite magnet which is magnetized in axial directions, and auxiliary ferrite magnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003970_ssrr50563.2020.9292576-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003970_ssrr50563.2020.9292576-Figure8-1.png", + "caption": "Fig. 8. Comparison of the grousers with different length and intervals", + "texts": [ + " Owing to this, the grousers may have rolled over the obstacle and level at the corner, improving the hook capability. Alternatively, their shape may have resulted in the detachment. L-shaped or T-shaped grousers may be effective in preventing these consequences. The length of the grousers of the first fabricated robot, which is not mentioned in this paper, were larger and the distance between them was smaller than those fabricated afterward. This made the tips of the grousers contacted each other when the track was bent to adapt to the obstacle (Fig. 8); therefore, the corners could not be gripped in the gap between the grousers, hampering the climbing. To avoid this issue, the dimensions of the grouser should be reduced, and the distance between them should be increased. The reduction of dimensions may, however, negatively impact the hooking capability and mechanical resistance of the grousers. In addition, if the intervals between the grousers are narrow, the robot will be mostly contacted by one grouser on the ground, as the MW-Track has only one sprocket, resulting in instability when moving on a flatland" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003813_ffe.13393-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003813_ffe.13393-Figure1-1.png", + "caption": "FIGURE 1 The three-dimensional geometry of wind turbine gearbox. A, Wind turbine gearbox; B, intermediate stage of wind turbine gearbox", + "texts": [ + " Considering the influences of inclusion size, inclusion position and residual stress, Section 3 presents the results of stress field near the inclusion and the fatigue life predictions based on Miner\u2019 theory. Finally, the conclusions are summarized in Section 4. As the main component of wind turbine driving system, the increasing gears with large transmission ratio could increase the speed of the wind turbine for the requirement of asynchronous generator. That is, the low-speed gear is served as the driving wheel, which transmits the low-speed torque from the blade side to the middle-speed gear shaft. As shown in Figure 1A, wind turbine gearbox can be divided into the low-speed stage, intermediate stage and high-speed stage. Specially, the threedimensional (3D) model of intermediate stage of wind turbine gearbox consisting of an involute helical gear shaft and the low-speed gear is given in Figure 1B. The geometry parameters of the gear shaft and low-speed gear are presented in Table 1. In the present study, the stress and fatigue life analysis are proposed for the specific intermediate gear shaft with oxide inclusion. For the intermediate gear shaft, it is found that some broken teeth occur at the time long before its design life (e.g. 20 years). The fracture surface of the failure gear shaft is presented in Figure 2. According to the experimental observations, it is considered that the fracture originates from the position 8 mm below the gear surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure6.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure6.4-1.png", + "caption": "Figure 6.4 A spherical joint.", + "texts": [ + "18) For the version shown on the right-hand side of Figure 6.3, the rotation sequence is ab rot[u\u20d7(ab) i = u\u20d7(bab) i ,\ud835\udf03ab] \u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2192bab rot[u\u20d7(bab) k = u\u20d7(ba) k ,\ud835\udf19ab] \u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2212\u2192ba (6.19) The corresponding orientation equation is C\u0302(ab,ba) = eu\u0303i\ud835\udf03ab eu\u0303k\ud835\udf19ab (6.20) Joints and Their Kinematic Characteristics 131 A spherical joint is also known as a globular joint or a ball-and-socket joint. Indeed, it consists of two kinematic elements that have the shapes of a ball and a socket as illustrated in Figure 6.4. Within the encasement of the socket, the ball can rotate freely about any arbitrary direction. Therefore, the mobility of a spherical joint is \ud835\udf07ab = 3, which is only due to the rotational freedom. Due to the encasement of the socket, the characteristic location equation is r(ab) ab,ba = 0 (6.21) As for the characteristic orientation equation, it can be written in the following general form in terms of three independent angular joint variables (\ud835\udf19ab, \ud835\udf03ab, \ud835\udf13ab). C\u0302(ab,ba) = en\u0303ab1\ud835\udf19ab en\u0303ab2\ud835\udf03ab en\u0303ab3\ud835\udf13ab (6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000380_chicc.2015.7259616-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000380_chicc.2015.7259616-Figure1-1.png", + "caption": "Fig. 1: Unknown input u of the studied system.", + "texts": [], + "surrounding_texts": [ + "By choosing N(\u03b4) = L(\u03b4)K(\u03b4) + G(\u03b4) and L(\u03b4) = Q(\u03b4)A(\u03b4)\u2212G(\u03b4)C(\u03b4) and , then we get\ne\u0307 = [Q(\u03b4)A(\u03b4)\u2212G(\u03b4)C(\u03b4)] e (9)\nAccording to Lemma 1, the second condition in Theorem 1 implies that Ol\u2217(\u03b4) defined in (8) is left unimodular over R[\u03b4]. In, it has been proven that if Ol\u2217(\u03b4) is left unimodular over R[\u03b4], then there exists a left unimodular matrix T (\u03b4) \u2208 R pl\u2217\u00d7n[\u03b4] such that\nT (\u03b4)Q(\u03b4)A(\u03b4)T\u22121 L (\u03b4) = A0 + F (\u03b4)C0 C(\u03b4)T\u22121 L (\u03b4) = C0\n(10)\nwhere F (\u03b4) = [ FT 1 (\u03b4), \u00b7 \u00b7 \u00b7 , FT l\u2217 (\u03b4) ]T and\nA0 = \u23a1 \u23a2\u23a2\u23a2\u23a3 0 Ip 0 \u00b7 \u00b7 \u00b7 0 ... ... ... . . . ...\n0 \u00b7 \u00b7 \u00b7 0 \u00b7 \u00b7 \u00b7 Ip 0 0 0 \u00b7 \u00b7 \u00b7 0\n\u23a4 \u23a5\u23a5\u23a5\u23a6 \u2208 R pl\u2217\u00d7pl\u2217\nC0 = [ Ip 0 \u00b7 \u00b7 \u00b7 0 ] \u2208 R p\u00d7pl\u2217\n(11)\nMoreover, the matrix T (\u03b4) is given by:{ T1(\u03b4) = C(\u03b4)\nTi+1(\u03b4) = Ti(\u03b4)Q(\u03b4)A(\u03b4)\u2212 Fi(\u03b4)C(\u03b4) (12)\nfor 1 \u2264 i \u2264 l\u2217 \u2212 1, with Fi(\u03b4) being determined through the following equations:\n[Fl\u2217(\u03b4), \u00b7 \u00b7 \u00b7 , F1(\u03b4)] = C(\u03b4) [Q(\u03b4)A(\u03b4)] l\u2217 [Ol\u2217(\u03b4)] \u22121 L\n(13)\nTherefore, if \u2203l\u2217 \u2208 N such that rankR[\u03b4]Ol\u2217(\u03b4) = n and InvS [Ol\u2217(\u03b4)] \u2282 R, then there exists T (\u03b4) such that:\nT (\u03b4) [Q(\u03b4)A(\u03b4)\u2212G(\u03b4)C(\u03b4)]T\u22121 L (\u03b4)\n= A0 + F (\u03b4)C0 \u2212 T (\u03b4)G(\u03b4)C0\nBy choosing G(\u03b4) = T\u22121 L (\u03b4) [F (\u03b4) +G0], we obtain\nQ(\u03b4)A(\u03b4)\u2212G(\u03b4)C(\u03b4) = T\u22121 L (\u03b4) [A0 \u2212G0C0]T (\u03b4)\nSince the pair (A0, C0) is observable, thus there exists a constant matrix G0 such that T\u22121\nL (\u03b4) [A0 \u2212G0C0]T (\u03b4) is Hurwitz independent of the time-delay, therefore Q(\u03b4)A(\u03b4) \u2212 G(\u03b4)C(\u03b4) in (9) is Hurwitz, and this implies that the observation error e converges to zero exponentially.\n5 Illustrative example\nConsider the following example:\nA(\u03b4) =\n\u23a1 \u23a3 \u03b42 1 \u03b4\n\u03b4 \u03b4 1 + \u03b4 1 \u03b4 \u03b42\n\u23a4 \u23a6 , B(\u03b4) = \u23a1 \u23a3 1\n1 \u03b4\n\u23a4 \u23a6\nand\nC(\u03b4) = [ 1 0 0 \u03b4 1 0 ] Since B(\u03b4) is already of full column rank over R[\u03b4], then V (\u03b4) = I, i.e. B(\u03b4) = B\u0304(\u03b4). Therefore we have\nC(\u03b4)B(\u03b4) =\n[ 1\n1 + \u03b4\n] and it can be checked that\nInvS\n[ C(\u03b4)B(\u03b4)\nB(\u03b4)\n] = InvS [ C(\u03b4)B(\u03b4) ] = {1}\nAccording to Lemma 2, one can find\nK(\u03b4) =\n\u23a1 \u23a3 1 0\n1 0 \u22121 1\n\u23a4 \u23a6\nand Q(\u03b4) = I\u2212K(\u03b4)C(\u03b4)\n=\n\u23a1 \u23a3 0 0 0\n\u22121 1 0 1\u2212 \u03b4 \u22121 1\n\u23a4 \u23a6\nsuch that Q(\u03b4)B(\u03b4) = 0. With this Q(\u03b4), we can find that there exists l\u2217 = 2, such that\nOl\u2217(\u03b4) = \u23a1 \u23a2\u23a2\u23a3\n1 0 0 \u03b4 1 0 0 0 0 \u03b4 \u2212 \u03b42 \u03b4 \u2212 1 1\n\u23a4 \u23a5\u23a5\u23a6\nwith rankR[\u03b4]Ol\u2217(\u03b4) = 3 and InvS [Ol\u2217(\u03b4)] = {1, 1, 1} \u2282 R. Therefore, all conditions of Theorem 1 are satisfied, and we can follow the proposed method to design an exponential observer.\nFor this, let us firstly determine T (\u03b4). According to (13),\nwe obtain\nF (\u03b4) = \u23a1 \u23a2\u23a2\u23a3\n0 0 0 \u03b4 \u2212 2 0 0\n1\u2212 \u03b43 0\n\u23a4 \u23a5\u23a5\u23a6\nwhich gives us, based on (12), the following T (\u03b4):\nT (\u03b4) = \u23a1 \u23a2\u23a2\u23a3\n1 0 0 \u03b4 1 0 0 0 0 3\u03b4 \u2212 2\u03b42 1 1\n\u23a4 \u23a5\u23a5\u23a6\nand its left inverse: T\u22121 L (\u03b4) =\n\u23a1 \u23a3 1 0 0 0\n\u2212\u03b4 1 0 0 2\u03b42 \u2212 2\u03b4 \u22121 0 1\n\u23a4 \u23a6.\nAccording to (10), we have\nA0 = \u23a1 \u23a2\u23a2\u23a3\n0 0 1 0 0 0 0 1 0 0 0 0 0 0 0 0\n\u23a4 \u23a5\u23a5\u23a6 , C0 = [ 1 0 0 0 0 1 0 0 ]\nThen we can freely choose G0 such that A0\u2212G0C0 is stable. In this example, we take\nG0 = \u23a1 \u23a2\u23a2\u23a3 1250 0 0 60\n3750 0 0 800\n\u23a4 \u23a5\u23a5\u23a6\nwhich sets the eigenvalue of A0 \u2212 G0C0 as (\u221275,\u221250,\u221240,\u221220). Finally we have\nG(\u03b4) = T\u22121 L (\u03b4) [F (\u03b4) +G0]\n=\n\u23a1 \u23a3 125 0\n\u2212125\u03b4 58 + \u03b4 \u2212\u03b43 + 250\u03b42 \u2212 250\u03b4 + 1 724\u2212 \u03b4\n\u23a4 \u23a6", + "L(\u03b4) = Q(\u03b4)A(\u03b4)\u2212G(\u03b4)C(\u03b4)\n=\n\u23a1 \u23a3 \u2212125 0 0\n68\u03b4 \u2212 2\u03b42 \u221259 1 \u2212493\u03b4 \u2212 248\u03b42 \u2212741 \u22121\n\u23a4 \u23a6\nN(\u03b4) = L(\u03b4)K(\u03b4) +G(\u03b4)\n=\n\u23a1 \u23a3 0 0\n\u22122\u03b42 \u2212 57\u03b4 \u2212 60 \u03b4 + 59 \u2212\u03b43 + 2\u03b42 \u2212 743\u03b4 \u2212 739 741\u2212 \u03b4\n\u23a4 \u23a6\nThe simulation results are depicted in the following figures.\n6 Conclusion\nThis paper considered a class of linear system with commensurate delay, involved in the model and in the input, which could be multiple. For such kind of systems, sufficient conditions have been given in order to guarantee the existence of a simpler Luenberger-like observer, involving only the past and actual values of the system output, which can exponentially estimate the states of the studied systems.\nReferences [1] F. J. Bejarano, T. Floquet, W. Perruquetti, and G. Zheng. Ob-\nservability and detectability of singular linear systems with unknown inputs. Automatica, 49(3):793\u2013800, 2013. [2] F. J. Bejarano and G. Zheng. Observability of linear systems\nwith commensurate delays and unknown inputs. Automatica, 50(8):2077\u20132083, 2014.\n[3] S. Bhattacharyya. Observer design for linear systems with\nunknown inputs. IEEE Transactions on Automatic Control, 23(3):483\u2013484, 1978. [4] D. Boutat, G. Zheng, J.-P. Barbot, and H. Hammouri. Ob-\nserver error linearization multi-output depending. In IEEE CDC, 2006. [5] M. Darouach. Linear functional observers for systems with\ndelays in state variables. IEEE Transactions on Automatic Control, 46(3):491\u2013496, 2001. [6] M. Darouach, M. Zasadzinski, and S. Xu. Full-order ob-\nservers for linear systems with unknown inputs. IEEE Transactions on Automatic Control, 39(3):606\u2013609, 1994. [7] E. Emre and P. Khargonekar. Regulation of split linear sys-\ntems over rings: Coecient-assignment and observers. IEEE Transactions on Automatic Control, 27(1):104\u2013113, 1982. [8] G. Hostetter and J. Meditch. Observing systems with un-\nmeasurable inputs. IEEE Transactions on Automatic Control, 18:307\u2013308, 1973. [9] M. Hou and P. Muller. Design of observers for linear sys-\ntems with unknown inputs. IEEE Transactions on Automatic Control, 37:871\u2013875, 1992. [10] M. Hou, M. H. Zitek, and R. Patton. An observer design for\nlinear time-delay systems. IEEE Transactions on Automatic Control, 47:121\u2013125, 2002. [11] A. Krener. (Adf,g), (adf,g) and locally (adf,g) invariant\nand controllability distributions. 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In 49th IEEE Conference on Decision and Control, pages 443\u2013448, Atlanta, GA, USA, December 2010. [18] G. Zheng, J.-P. Barbot, D. Boutat, T. Floquet, and J.-P.\nRichard. On observation of time-delay systems with unknown inputs. IEEE Transactions on Automatic Control, 56(8):1973\u20131978, 2011. [19] G. Zheng and D. Boutat. Synchronisation of chaotic systems\nvia reduced observers. IET Control Theory & Applications, 5(2):308\u2013314, 2011. [20] G. Zheng, D. Boutat, and J. Barbot. A single output depen-\ndent observability normal form. SIAM Journal on Control and Optimization, 46(6):2242\u20132255, 2007. [21] G. Zheng, D. Boutat, and J.-P. Barbot. Output dependent\nobservability linear normal form. In IEEE CDC-ECC, 2005. [22] G. Zheng, D. Boutat, and J.-P. Barbot. Multi-output depen-\ndent observability normal form. Nonlinear Analysis: Theory, Methods & Applications, 70(70):404\u2013418, 2009. [23] G. Zheng, Y. Orlov, W. Perruquetti, and J.-P. Richard. Fi-\nnite time observer-based control of linear impulsive systems with persistently acting impact. In IFAC World Congress, volume 18, pages 2442\u20132447, 2011." + ] + }, + { + "image_filename": "designv11_34_0001131_icrom.2013.6510116-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001131_icrom.2013.6510116-Figure4-1.png", + "caption": "Figure 4 -Double inverted pendulum model in ADAMS", + "texts": [], + "surrounding_texts": [ + "The dynamics and kinematics model of a two wheel mobile manipulator robot with a reaction wheel is described in this section. For the considered robot which is shown in Figures (1-2), its parameters specifications are given in Table (1). The base of the mobile manipulator is assumed as a passive joint. Therefore, the manipulator part consists of three links and actuators in 3D workspace. The reaction wheel is placed on the first link which has one more degree of freedom rather than this one. The dynamics equations of motion of a mobile manipulator are described as: '( = M(q)q + H(q,q) + G(q) (1) where q=[ql'q2'\" .. ,qtER7 is the vector of generalized coordinates and =[\" 2,0, 4' 5' 6' lER7 is the input generalized torque and M(q) ER7x7, H(q, q) ER7xI, G( q) ER7X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively. The considered two WMM robot with a reaction wheel required complex dynamics modelling as a result of under actuated system. The passive joint causes that the balancing challenge of this robot in the XOY plane is more important than other its positions. Thus, Double Inverted Pendulum Model (DIPM) is utilized to simplify the dynamic analysis for the balancing control, [18]. The effectiveness of this simplified model is demonstrated for the dynamic locomotion of the highly nonlinear and complex system, [19]. Tablel - Parameters of the two WMM with a reaction wheel o XYZ OI_XIYIZI Q,&q2 Q3 Q;(i=4,5,6) q, L 1,(i= I ,2,3) 14 Word coordinate frame Mobile manipulator coordinate frame Rotation angles of wheels Inclination angle of the passive joint Joint angles of links Angular position of the reaction wheel Radius of wheels Distance between wheels Length of links Distance of the reaction wheel centre to first joint B. Dynamics Modelling of Double Inverted Pendulum It is possible to model two WMM as a virtual double inverted pendulum model. In this model, the components of second and third of manipulator in XOY plane are considered as the virtual second link of double inverted pendulum model. This simple model is used as the stabilizing control part. This model is shown in Figure (3) and parameters of these models is stated in Table (2). Table2 - Parameters of DIPM ilw Rotation angle of wheels ill Joint angle of the first link ile Joint angle of the second link il2 Angular position of the reaction wheel 11 Length of the first link Ie Length of the second I ink 12 Distance of reaction wheel centre to first joint xme X position of the me in 01 coordinate frame Yme Y position of the me in 01 coordinate frame Xl X component of the first link Length Y1 Y component of the first link Length me Equivalent mass on the second link m1 Mass of the first link m2 Mass of the reaction wheel mw Mass of wheels c,(i=2,e) Coefficient of friction a1 Length of the CoG on the first link i1 Moment inertia of link I around the ZI axis i2 Moment inertia of the reaction wheel around the zraxis Calculation of positions of the virtual link xmeand Yme' and its length Ie' joint angleqe and the mass meare as following: (2) (3) (4) (5) The dynamics equations of motion of this model are obtained by Euler-Lagrange equation as: (6) whereq = [qw, q1' q2' qeF ER4 is the vector of generalized coordinates and r = [fw, f1' f2' feF ER4 is the input generalized torque for the virtual double inverted pendulum modeI.M(q) ER4X4,A(q, q) ER4XI,G(q) ER4X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively which are calculated as: M = [\ufffd\ufffd\ufffd M31 M41 where: Mll = (mw + m1 + mz + me)rZM12 = M21 = (a1m1 + me (11 + Ie COS(qe)) + 11mz)rcOS(q1) + mwrZM14 = M41 = melerCOS(q1 + qe) Mzz = i1 + iz + a1 Z m1 + 11 z mz + 11 z me + Ie z me cosz qe + 211 leme cos qe + mWrZM23 = M32 = M33 = izMz4 = Ie z me cosz qe + 11 leme cos qe M13 = M31 = M34 , Z = M43 = OM44 = Ie meH1 = -fiI(mer(ft1 sin(i'iI) (\\1 + Ie COS(qe)) + leqe sin(cl!+qe)) + 11mZft1rsin(Q1) + a1 m1 ib r Sin(Q1)) - merqe(leqe Sin(q1 + qe) + leq1 COS(q1) sin(qe))Hz = -2m.leq1 (feqe COS(qe) sin(qe) + 11 qe sin(qe)) + meleqe (qwr COS(q1) COS(qe) - qe sin(qe) (\\1 + 21e cos(qe))) H3 = cZqZH4 = -me Ie rih qw COS(qe) sin(qe) + m.leq1 z sin(qe) (11 + Ie COS(qe)) + CeqeG1 = G3 = OGz = -a1 m1 Sin(q1) - 11 sin q1 (mz + me) - Ie me (cos q1 sin qe + sin q1 cos qe)G4 = -Ieme sin(q1 + qe) C. Dynamic Verification/or DIPM ADAMS is the most widely used multi-body kinematics and dynamics analysis software in the word. Also, Adams helps engineers to study the kinematics and dynamics of moving parts, how loads and forces are distributed throughout mechanical systems, and to improve and optimize the performance of their products. The extracted dynamics equations of motion of the virtual double inverted pendulum with a reaction wheel are verified by the ADAMS model which the 3D sketch of DIPM is shown in Figure (4).We insert some similar torque as input to these two models and compare the reaction of model joints with together. This torques is shown in Figure (5) which insert to the reaction wheel. Moreover, the obtained result of the model response is shown in Figure (6).These curves show that the dynamics equations of motion of the virtual double inverted pendulum are verified. So, we can use these equations in the control algorithm to realize the dynamic stability. In addition, the parameters specifications of this system in this verification routine are expressed in Table (3). 0.4 0.3 0.2 E 0.1 ;;. ! -0.1 -0.2 -0.3 I I I I I I I I I ___ 1 ____ 1___ -.l ___ .l ___ L ___ 1 __ _ 1___ -.J ___ \ufffd __ _ I I I I -- \ufffd --- 1 --- T --- \ufffd -- -0.4 - - -1- - - -1 - - - -+ - - - + - - - t- - - - 1- - - -1- - - ---j I I I I I I I I -0.5 OL--- :0c'c .5=- -L.- 1\ufffd.5=- -\ufffd2 --2cc'.C:-5-\ufffd3---=-3.L5 --':-\ufffd4\"'.5-\ufffd 5 Time (8) Since, the two WMM contains a passive joint, it requires an active controller to stabilize the passive joint and to control its stability. Related to ZMP criterion,[20], this robot have just one stable natural position that the moment of the gravity has no effect on the body, in other words, gravity force direction of COG crosses the mobile wheel axis. If we control the position of the COG to this point, we can able to satisfy the robot stability. In this section, a PID controller is proposed to control the motion of the reaction wheel to achieve dynamic stability. The supervising controller identifies the position of the COG and tunes the set point of PID controller. At the same time, the PID controller moves the reaction wheel to reach the stable natural position. Also, the control block diagram of this controller is shown in Figure (7). The supervising controller finds the COG position without any information about robot links and the related body. So, as shown in Figure (7), the input parameters of this block are the reaction wheel angular velocity and its acceleration. Therefore, it can just find the position of the COG to the right or the left from the normal position of the first link (qi=O) with this information. For an example, in Figure (2), if the reaction wheel angular velocity and its acceleration are positive, it is clear that the position of COG are in the left side from the normal position of the first link (qi=O).SO, the supervising controller changes the PID set point to the negative value. Thus, this can use many type of controller for this block such as PI, PID, Fuzzy or GA controller. In this paper we used the PI controller for this block. IV. SIMULATION RESULTS AND DISCUSSION The validation of the proposed control strategy is demonstrated by simulation results. This simulation runs at MA TLAB Simulink Toolbox for the assumed robotic system which the parameters specifications are expressed in Table(3). In addition, the controller gains are expressed in Table (4). Table4--Controller gains amount Kp Coefficient of P action 3.916 KI Coefficient of I action 1 KD Coefficient of 0 action 0.0624 Ksd Coefficient of q3 parameter -6e-3 Ksp Coefficient of q3 parameter 0.0209 This type of robot is dynamically stable and this stability is attained using a reaction wheel. Here, we consider the three bench marks and run the simulations. In the first and second one, the robot start from an initial position and make itself stable and, in third one, the robot run in the different initial position and pass the variable acceleration. The initial conditions of these case studies are respectively expressed in Table (5). Moreover, the accelerate variation curve of the third case is shown on Figure (8). Table 5 - Initial conditions for simulations Casel Casell CaselII q qp i'i2 Joint angle of the first link (deg) -30 -30 30 Joint angle of the second link (deg) -60 -120 90 Angular position of the reaction wheel (deg) 0 0 0 1.5 - - - - - I - -----1- - - - - - 1 - - --- 1 ----- - - - - - ---1 - -----1- - - - - - r- - ---- _____ --J ______ 1 ______ L ____ _ ........ 0.5 1 .\ufffd \ufffd ] -0.5 I I I ----- T -----1-----l------------ \ufffd -----r----- \ufffd----- \ufffd------------ \ufffd.5 -----\ufffd----- 4----- \ufffd------------ -20\ufffd------\ufffd,0\ufffd----\ufffd2\ufffd0 ----\ufffd3\ufffd 0------\ufffd4 0\ufffd----\ufffd 5\ufffd 0------\ufffd6 0 Time(s) Figure 8 - Acceleration variation during the operation As shown from simulation results in Figures (9-17), the time history of the robot motion is presented in Figures (9- 11) for the cases to move and reach to the stable position. Moreover, the position of the <11 must be varied until the CoG reaches to the appropriate position Figures (12-14). In this position, the robot is stable and there is no moment appears to change this style. Also, the PID modifies the set point value to satisfy the stabilization. Figures (15-17) are shown the position of the reaction wheel, during the robot reaches to stable position. These results show that the reaction wheel movement makes the internal moment that it can control the passive joint. In this control strategy, the balancing is not achieved by the robot movement and only the reaction wheel is used to stabilize. This method can also improve the dexterity of the robot motion and increases the robot controllability." + ] + }, + { + "image_filename": "designv11_34_0002121_tcst.2016.2572165-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002121_tcst.2016.2572165-Figure3-1.png", + "caption": "Fig. 3. Definition of angle of attack \u03b1.", + "texts": [ + " (3) The relationship between angular velocity vector in body-fixed frame \u03c9 = [\u03c9x , \u03c9y, \u03c9z ]T and the vector of Euler angle rates \u0307 = [\u03c6\u0307, \u03b8\u0307 , \u03c8\u0307]T is given by \u03c9 = B\u0307 (4) where B = \u23a1 \u23a3 1 0 \u2212 sin \u03b8 0 cos\u03c6 sin \u03c6 cos \u03b8 0 \u2212 sin \u03c6 cos\u03c6 cos \u03b8 \u23a4 \u23a6. (5) Matrix B becomes singular when \u03b8 = \u00b1\u03c0/2, and thus, we restrict the inverse transformation to \u2212\u03c0/2 < \u03b8 < \u03c0/2. Let us consider an unstable cylindrical-shaped rigid body. We assign the body-fixed frame at the CM, such that the x-axis is along the axis of the cylinder, and Y and Z axes define the cylinder circular cross-sectional plane. The angle of attack in 3-D, \u03b1, is the angle between body velocity vector, v, and the body x-axis as shown in Fig. 3. The ideal flight configuration is such that the body axial axis (x-axis) is aligned with its CM velocity vector, resulting in zero angle of attack. However, if the body velocity vector, v, and its x-axis are misaligned, \u03b1 will have a nonzero value and the two vectors form a plane, which can be identified by its normal, u\u03b1 . Thus, u\u03b1 is a unit vector normal to the plane containing the x-axis and v and defines the direction of the rate of change of the angle of attack \u03b1\u0307 u\u03b1 = i \u00d7 v ||i \u00d7 v|| (6) where i represents a unit vector in the body x-axis direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000123_cdc40024.2019.9029665-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000123_cdc40024.2019.9029665-Figure7-1.png", + "caption": "Fig. 7. (a) Experimental configuration for the real-time THz imaging system using THz camera. (b) Measured beam profile of THz wave without sample. (c) Measured transmission image of a dry leaf. (d) Transmission image of the dry leaf corrected for the non-uniform beam profile.", + "texts": [ + " 6(a) and (b) shows the beam profiles taken with the original THz camera and the sensitivity-improved THz camera, respectively. It is clearly seen that the SNR is increased with the sensitivity-improved THz camera. In particular, SNRs in the original THz camera and the sensitivity-improved THz camera were calculated to be 80 and 695, respectively, from the line profiles shown in Fig. 6. This indicates that an SNR improvement of one order of magnitude was achieved. We demonstrated THz transmission imaging of a dry leaf taken with the sensitivity-improved THz camera at the video rate. The results are shown in Fig. 7 and Video 1. To obtain these images, we used a THz wave generated from the LN crystal, the THz camera, and a pair of THz lenses made of Tsurupica (focal lengths of the lenses were 50 and 30 mm). The experimental configuration is shown in Fig. 7(a). We first obtained the reference beam profile of a THz wave without the sample, as shown in Fig. 7(b), and then acquired the dry leaf image, as shown in Fig. 7(c). The THz transmittance image of the dry leaf was obtained by dividing the signal value acquired in Fig. 7(c) by that acquired in Fig. 7(b) for every pixel, as shown in Fig. 7(d). The obtained movie is shown in the supplement. Video-rate THz transmission images were obtained with the sensitivity-improved THz camera. Clear THz images without interference fringes were obtained because we used THz pulses with broadband spectrum. VI. CONCLUSION We increased the sensitivity of a THz camera in the frequency range below 1 THz by increasing the effective cavity length in the micro-bolometer. We also developed a new method for evaluating the frequency dependence of THz camera sensitivity, using a broadband THz wave source and THz bandpass filters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002656_icamechs.2016.7813415-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002656_icamechs.2016.7813415-Figure15-1.png", + "caption": "Fig. 15. Results of the moving experiment at a slope(spike wheels)", + "texts": [ + " The experimental results are presented in Fig. 14. As shown in the figure, the distance of the system that we don't control was 1[m], slipping distance was 270[mm]. On the other hand, the distance was 2.5[m] and 180[mm] by the system that it controlled. Therefore, we check that the proposed control method is effective at a slope. Then, we confirmed about the effect of the spike wheels. We also estimate this experiment by the distance about the effect of the spike.The experimental results are shown in Fig. 15. Figure 15 shows that slipping is prevented by spike substantially. Finally, the safety of a trimmer-type edge without a cover is also studied. We confirm the safety by making the experiment which makes our system collide with a person intentionally. As shown in Fig. 16, the trimmer-type edge is safe. Even if the mowing edge collides with a person, it is safe because the fixed edge is in contact with a person. The results of these experiments showed that the system was safe to operate. VI. CONCLUSIONS In this study, we proposed and developed a trimmer-type mowing system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001301_s11665-015-1511-4-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001301_s11665-015-1511-4-Figure11-1.png", + "caption": "Fig. 11 Temperature history ( C) of transverse cross section of weld calculated using cross section information given in Table 2 for transformation boundary, where Dt = Dl/V, Dl = (12.7/60) mm, and V = 4.167 mm/s (Weld 2)", + "texts": [], + "surrounding_texts": [ + "Although the calculations presented here demonstrate certain aspects of the methodology, there still remain further investigations concerning its practical implementation, which would be in terms of algorithm development. It is therefore important to examine aspects of the inverse analysis methodology that are relevant to its further development and refinement. The general procedure for inverse thermal analysis of welds as described in this study includes interpolation between constrained isothermal boundaries, e.g., TTB and TM. A specific procedure for interpolation, however, has not been considered. Specific procedures for interpolation between constrained Journal of Materials Engineering and Performance surfaces within a temperature field are therefore an issue for algorithm development. Accordingly, further investigation is needed to determine a general and optimal procedure for interpolation between constrained isothermal surfaces in three dimensions. Toward this goal, the formal mathematical foundation of interpolation between constrained boundaries and various paths for linear interpolation is discussed. Formally, a solution to the steady-state advection-diffusion equation V @T @x \u00bc r \u00f0jrT\u00de; \u00f0Eq 6\u00de is to be determined within the region of the temperature field bounded by isothermal boundaries at TTB and TM, and surface and symmetry plane boundaries, which are nonconducting. This region of the temperature field is such that Eq 6 can be solved sequentially within one-, two-, and three-dimensional subdomains using direct model-based numerical procedures. These subdomains and the associated partitioning of Eq 6 are shown in Fig. 14 for the calculated temperature field of Weld 2. Following an inverse analysis approach, however, a solution to Eq 6 can be calculated Journal of Materials Engineering and Performance using interpolation procedures, which can be applied owing to the close proximity of boundary conditions, in contrast to more complex direct model-based procedures. The numerical solution of Eq 6, for specified boundary conditions, which is by application of various numerical algorithms (e.g., methods based on finite differences and volumes), is of the general form Journal of Materials Engineering and Performance Tn\u00fe1 p \u00bc P6 k\u00bc1 wkTn k P6 k\u00bc1 wk ; \u00f0Eq 7\u00de where the weight coefficients wk and iteration sequences, indexed by n, are specified according to a given solution algorithm. The grid indices k specify nearest neighbor grid locations relative to the grid location specified by index p. The general form of Eq 7, with constraints conditions specified according to Eq 3b, is that of the discrete interpolation operator. Accordingly, an important property following from this form is that for temperature field regions between boundaries, which are closed and in close proximity, regardless of the type of numerical procedure, the numerical solution of Eq 6 is equivalent to interpolation between boundary values. A heuristic discussion of this property, which is associated with the transition of numerical solvers for parabolic equations to those for elliptic equations, according to boundary conditions, is given in reference (Ref 24) with respect to finite volume discretization. Examination of the temperature field boundary conditions associated with both workpiece geometry and constrained isothermal surfaces TM and TTB establishes various paths for linear interpolation (see Fig. 14). Following these paths, interpolation between constrained surfaces can in principle be partitioned into a series of one-, two-, and three-dimensional interpolations. Referring to Fig. 14, one observes that boundary conditions on the calculated temperature field are such that interpolation can be applied in one dimension along the symmetry line at the top surface, in two dimensions over the top surface and symmetry planes, and in three dimensions over the closed domain bounded by isothermal surfaces TM and TTB, and the top surface and symmetry planes. As discussed previously (Ref 2), the embedding of constrained isothermal surfaces, which are constructed according to experimental measurements, within the calculated temperature field, compensates for errors due to estimated values of the diffusivity function, which are in fact a function of temperature. This follows in that embedded isothermal surfaces, which are sufficiently distributed volumetrically, represent implicitly the dependence of diffusivity on temperature, as well as advective influences within the melt pool. Multiple constrained isothermal surfaces can be constructed using thermocouple measurements following the same procedure applied in this study. Journal of Materials Engineering and Performance Parametric temperature histories obtained by inversion can be adopted for construction of databases, which in turn can provide for optimal inverse analysis of welding processes spanning a wide range of process parameters. The concept of constructing such databases is well posed in that, although there exist many different types of welding processes and their potential modifications, the physical characteristics of all types of welds are within a limited range of variation. That is to say, for all types of welding processes, the range of variations in solidification boundary shapes (and associated isothermal surfaces), and in weld thickness, is finite. In addition, for all metals and their alloys, the range in variation of material properties is also finite. It follows that a database of parametric temperature histories can be constructed that spans a wide range of welding processes, as well as their potential modifications. The availability of a parametric temperature history database can provide both parameter estimates for inverse analysis and correlation with documented weld analyses, both experimental and computational. The relationship between such a database and other components of the general framework for inverse thermal analysis using constraints is shown in Fig. 15. Referring to Fig. 15, it is seen that the results of both laboratory and computational experiments provide constraint conditions for inverse thermal analysis, while basic theory (heat transfer due to advection and diffusion) provides formulations of parametric models, i.e., different types of basis functions. In addition, the general framework shown in Fig. 15, which describes the use of constraint conditions obtained from computational experiments, establishes a well-defined relationship between direct and inverse modeling. An elucidation of this relationship is of significance in that many times direct and inverse models are compared unfairly. A discussion of such comparisons is beyond the scope of this study. It should be noted, however, that because of the interrelationship between direct and inverse models, it should be difficult to establish common criteria for their general comparison. In general, both direct and inverse models have aspects that contribute to computational cost, especially with respect to optimum use of cutting edge computational resources. As discussed above, the computational cost of direct models follows naturally from detailed mathematical representation of underlying physical processes. In contrast, the computational cost of inverse models follows from the ill-posedness of inverse problems, whose solutions are not unique. In particular, the inverse heat conduction problem (IHCP) is generally ill-posed, and thus methods of problem regularization are required, which tend to be iterative and computationally intensive (Ref 16). The \u2018\u2018inverse weld analysis problem\u2019\u2019 as specified in this study, although related to the IHCP, is well posed and not computationally intensive, as demonstrated by this and other studies (Ref 1, 2). References (Ref 16, 25) present discussion of illposedness of inverse problems associated with the IHCP and engineering applications, in general. The inclusion of a parametric temperature history database within the general inverse analysis framework (see Fig. 15) provides, in addition to parameter estimates, correlations with other type of weld analyses, such as microstructural analysis, as well as with weld processes and associated process parameters that are both feasible and optimal, which are documented in the literature. Finally, the present study applies a parametric model that is in terms of numerical-analytical basis functions, whose parameters are conveniently adjustable with respect to multiple constraint conditions. This parametric model is an inverse model formulation at its foundation. Direct model formulations, however, can also be applied parametrically for inverse analysis. For example, references (Ref 11, 26-28) present inverse analyses using direct models, and associated numerical simulations, for the determination of unknown variables. An issue to be addressed for inverse analysis using direct models is the feasibility of parameter adjustment with respect to multiple constraint conditions." + ] + }, + { + "image_filename": "designv11_34_0002485_j.triboint.2016.10.041-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002485_j.triboint.2016.10.041-Figure4-1.png", + "caption": "Fig. 4. Rotating part: case and assembly.", + "texts": [ + " The circuit represented in Fig. 2 must be integrated in the test rig by mounting it on the rotating ring. Thus, it must be in-built with the rotating shaft. For this purpose, a proper PVC collar has been designed and produced by 3D printing. The collar has been covered with an aluminum case, with the aim of covering but also containing the circuital modules, even during the rotation, to prevent collapses due to the centrifuge force. The assembly of the collar and its case has been keyed on the shaft (Fig. 4). Once the circuit was placed into the collar and covered by the aluminum case, the thermocouple probe has been placed in contact with the inner portion of the rotating ring through a hole, drilled in the rotating ring housing and in the rotating counter-face of the seal, as shown in Fig. 5. The thermocouple probe has been introduced into the hole and, taking into account the high centrifuge forces during the rotation, a reliable contact between the probe and the ring has been assured. Another thermocouple is connected to the stationary ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001689_s11071-015-2356-y-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001689_s11071-015-2356-y-Figure5-1.png", + "caption": "Fig. 5 Demonstration of a unstable point with respect to slipping, the trajectory remains in Case NS permanently. Left panel effect of perturbation in time. Right panel projection of the trajectories onto the orthogonal space", + "texts": [ + " For the stationary solution (46) of Case NR, \u03b21 = 0, \u03b22 = \u03c0 and we obtain \u03b3 (x\u2217, \u03b21) = \u2212 j+1 r j ( \u03b7R\u03c92 0 \u2212 g ) < 0, \u03b3 (x\u2217, \u03b22) = \u2212 j+1 r j ( \u03b7R\u03c92 0 + g ) < 0, (60) that is, (46) is stable with respect to slipping if \u03c90 > \u221a g \u03b7R . (61) In this case, the dynamics extinguishes the effect of the perturbation, and after awhile, the ball returns to the stationary solution of Case NR (see Fig. 4). However, if \u03c90 < \u221a g/(\u03b7R), then the stationary solution is unstable with respect to slipping and the dynamics repels the trajectories away from the stationary solution of Case NR (see Fig. 5). If only neutral stability is required then instead of (61), we get the same result as in (46), which was obtained from the requirement (13). By direct calculation from (27) and (56), one can check that this agreement is valid not only for the stationary solution but also for any point in XNR. The reasons behind this coincidence can be explained from the properties of the simple Coulomb friction model (see [12], p. 85 and p. 139). In this section, we focus on the stationary solution (32) of Case RR" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003782_s12555-019-0436-3-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003782_s12555-019-0436-3-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the CV.", + "texts": [ + " Thirdly, an EKF is used to get the best position estimation of the CV by combining the encoder positioning results and landmark positions obtained from laser sensor Lidar. Fourthly, a MIMO robust servo controller for tracking the desired trajectory is designed by using the LSID operator. Finally, to verify the effectiveness of the proposed SLAM algorithm and MIMO robust servo controller, the experimental results are presented. These experimental results of the proposed MIMO robust servo controller are compared with those of the backstepping control method proposed by Pratama in [4]. The schematic modeling of a Caterpillar Vehicle (CV) is shown in Fig. 1. In this Fig. 1, [XA, YA] T is the position vector of the CV in a global coordinate frame OXY with orientation angle \u03b8A, VA is a linear velocity, and \u03c9A is the angular velocity of the CV. This CV has two driving standard wheels W1 and W2 with radius r, vR and vL are the right and left wheel linear velocities, respectively, \u03c6\u0307R and \u03c6\u0307L are the right and left wheel angular velocities, respectively, C is the origin of the coordinate frame Cxy, b is a distance between wheels and Cx axis of the CV, D is a mass center of the CV, d is the distance between C and D, and W3 and W4 are passive wheels of the CV. The kinematic equation of the CV as shown in Fig. 1 can be expressed as follows: x\u0307v =H\u03b7 , (1) xv = XA YA \u03b8A , H = cos\u03b8A 0 sin\u03b8A 0 0 1 , \u03b7 = [ VA \u03c9A ] , xv = \u222b x\u0307vdt, (2) where xv is a posture vector of the CV in global coordinate OXY and \u03b7 is a velocity vector of the CV. Relation of \u03b7 with the right and left wheel angular velocities (\u03c6\u0307R, \u03c6\u0307L) can be expressed as \u03b7 = T \u03a6\u0307 or \u03a6\u0307 = T \u22121 \u03b7 , (3) where T = [ r 2 r 2 r 2b \u2212r 2b ] , \u03a6\u0307 = [ \u03c6\u0307R \u03c6\u0307L ] . By substituting (3) into (1), the kinematic equation of CV can be expressed as follows: x\u0307v = S\u03a6\u0307 for S =HT " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003829_s10846-020-01290-1-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003829_s10846-020-01290-1-Figure3-1.png", + "caption": "Fig. 3 The sketch of neighborhood-sampling method. The red and green points are the sample points which receive more and less attention, respectively", + "texts": [ + " After the needle tip localization, all the pixels in needle region on Is are set to zero because the needle region information is not used in the following procedures. The neighborhood-sampling module is used to extract collision aware information. In general, the desired movement action is depended on the neighborhood state of current position and the goal position. And the collision-related information around the current position should be extracted effectively. Therefore, we propose a neighborhood-sampling method based on the image segmentation result. The sampling method is illustrated in Fig. 3. The receptive field of needle tip is a circle area on the segmentation map. The circle\u2019s radius is rf . A grid of discrete points {pt i}(i = 0, 1, 2, . . . , N), are sampled on Is within the circle and the sampling interval distance is dr . The index values at these sampling points are concatenated as a vector, i.e. the environment state se, whose element values are 0 or 1. The dimension of se is N . se = [Is(pt1), Is(pt2), . . . , Is(ptN )]T (1) Furthermore, the importance of the information at different sampling positions is not the same for the movement inference. Therefore, we embed the attention mechanism in the sampling module, which uses the fully connected layers and softmax function to infer a weight vector, so that the index value of each sampling point is rescaled. Intuitively, the sampling points near or within the obstacle region should be given larger weights in multiplication. The sampling points far from the obstacle region could be suppressed by given them smaller weights in multiplication. As is shown in Fig. 3, the attention mechanism is expected to assign different attentions to different parts of reception field, according to the whole environment state. The attention mechanism is realized by two fully connected layers, which transfers the environment state se to wa, namely, wa = \u03c3(W2(\u03b4(W1se))) (2) where W1 and W2 are the weights of two fully connected layers. \u03b4(\u00b7) and \u03c3(\u00b7) are the ReLU activation function and softmax output function, respectively, as given by, \u03c3(xi ) = exi \u2211N i=1 exi , \u03b4(x) = { x, x \u2265 0 0, x < 0 (3) Finally, the rescaled environment state sf is the elementwise multiplication of se and wa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000133_978-0-85729-588-0-Figure17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000133_978-0-85729-588-0-Figure17-1.png", + "caption": "Fig. 17 Final design of the RL2 Hand", + "texts": [ + " The extraction system of the RL2 Hand fingers keeps almost intact its functioning philosophy because the results obtained at the RL1 Hand has been successful. In this particular case, the design has been only focused on the energetic performance of the energy transformer. Pulleys have been located in all the tendons way avoiding them to pass over surfaces that can create unnecessary friction. In Fig. 16, it is possible to see the changes done for this last design and a detail of the pieces that compound it. The final result of the design is shown in Fig. 17. Robotic Hands 37 Ambrose R, Aldridge H, Askew R, Burridge R, Bluethmann W, Diftler M, Lovchik C, Magruder D, Rehnmark F (2000) Robonaut: NASA\u2019s space humanoid. IEEE Intell Syst Appl 15 Balaguer C, Gim\u00e9nez A, Jard\u00f3n A, Correal R, Cabas R, Staroverov P (2003) Light weight autonomous robot for elderly and disabled persons\u2019 service. Int Conf Field Serv Rob Balaguer C, Gim\u00e9nez A, Jard\u00f3n A, Cabas R, Correal R (2005) Live experimentation of the service robot applications elderly people care in home environments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003764_j.prostr.2020.10.069-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003764_j.prostr.2020.10.069-Figure1-1.png", + "caption": "Fig. 1. Symmetrical printing strategy utilised to manufacture four-sided hollow boxes. (The arrows indicate the print nozzle movement along the toolpath).", + "texts": [ + "25 g/mm using a RepRap X400 3D-printing system. Specimens were produced as hollow four-sided boxes with walls formed by individual extruded filaments. Geometry was defined by direct custom GCODE created with scripts and in-house software (FullControl GCODE Designer \u2013 contact the corresponding author for details). This enabled precise and accurate nozzle positioning throughout the deposition process as well as explicit control of print speed, extrusion volume and manufacturing sequence (unidirectional deposition) as shown in Fig. 1. dogbone specimens for tensile testing were developed at the scale of individual extruded filaments by controlling the volume of extrusion at different stages of the toolpath sequence (Fig. 2), made possible by the direct GCODE development. Two four-sided box types were generated, differentiated by the orientation of their dogbone form relative to the build orientation (Fig. 2). This enabled generation of specimens loaded in the F direction (longitudinally) (Fig. 2 (a)) and Z direction (transverse - normal to filaments) (Fig", + "5 mm in the gauge region, incrementally widening to a shoulder with thickness of approximately 2x the gauge thickness (controlled by extrusion volume). The boxes were deposited from a heated nozzle at 210 \u00b0C and platform temperature of 60 \u00b0C. Printhead speed was constant at 1000 mm-1. Custom GCODE ensured the nozzle position and deposition sequence was precisely controlled at all stages. Thus, any direct thermal variation across the geometry was prevented by maintaining a symmetrical toolpath (as shown in Fig. 1) and a constant print speed. Every region of the part geometry was subjected to the same cooling times and shared the same thermal history. Fig. 1. Symmetrical printing strategy utilised to manufacture four-sided hollow boxes. (The arrows indicate the print nozzle movement along the toolpath). Specimens were prepared by the following sequence: (i) The four-sided hollow boxes were cut at each of the corners with a razor blade mounted in a specially made tool, resulting in four equal walls. (ii) Using a second specially designed tool, the walls were cut using an array of seven razor blades spaced 5 mm apart from each another. A hydraulic press employed to compress the blades into the walls ensured even and consistent pressure and a clean cut", + "25 g/mm using a RepRap X400 3D-printing system. Specimens were produced as hollow four-sided boxes with walls formed by individual extruded filaments. Geometry was defined by direct custom GCODE created with scripts and in-house software (FullControl GCODE Designer \u2013 contact the corresponding author for details). This enabled precise and accurate nozzle positioning throughout the deposition process as well as explicit control of print speed, extrusion volume and manufacturing sequence (unidirectional deposition) as shown in Fig. 1. dogbone specimens for tensile testing were developed at the scale of individual extruded filaments by controlling the volume of extrusion at different stages of the toolpath sequence (Fig. 2), made possible by the direct GCODE development. Two four-sided box types were generated, differentiated by the orientation of their dogbone form relative to the build orientation (Fig. 2). This enabled generation of specimens loaded in the F direction (longitudinally) (Fig. 2 (a)) and Z direction (transverse - normal to filaments) (Fig", + "5 mm in the gauge region, incrementally widening to a shoulder with thickness of approximately 2x the gauge thickness (controlled by extrusion volume). The boxes were deposited from a heated nozzle at 210 \u00b0C and platform temperature of 60 \u00b0C. Printhead speed was constant at 1000 mm-1. Custom GCODE ensured the nozzle position and deposition sequence was precisely controlled at all stages. Thus, any direct thermal variation across the geometry was prevented by maintaining a symmetrical toolpath (as shown in Fig. 1) and a constant print speed. Every region of the part geometry was subjected to the same cooling times and shared the same thermal history. Specimens were prepared by the following sequence: (i) The four-sided hollow boxes were cut at each of the corners with a razor blade mounted in a specially made tool, resulting in four equal walls. (ii) Using a second specially designed tool, the walls were cut using an array of seven razor blades spaced 5 mm apart from each another. A hydraulic press employed to compress the blades into the walls ensured even and consistent pressure and a clean cut" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001689_s11071-015-2356-y-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001689_s11071-015-2356-y-Figure1-1.png", + "caption": "Fig. 1 Sketch of the mechanical model. The swirling flow of the fluid makes the ball travel round along the edge of the vessel", + "texts": [ + "eywords Dual-point contact \u00b7 Flowmeter \u00b7 Nonsmooth dynamics \u00b7 Discontinuity-induced bifurcations One of the basic principles of flow rate measurement is to place a solid body into the fluid and to let it bemoved M. Antali (B) \u00b7 G. Stepan Department of Applied Mechanics, Budapest University of Technology and Economics, Budapest 1521, Hungary e-mail: antali@mm.bme.hu G. Stepan HAS-BME Research Group on Dynamics of Machines and Vehicles, Hungarian Academy of Sciences, Budapest 1521, Hungary e-mail: stepan@mm.bme.hu or rotated by the fluid flow. One possible concept of this idea can be seen in Fig. 1, which can be found in several accepted patents from the last decades. A metal ball is placed into an axisymmetric vessel, in which swirling flow is created by blades or other geometric solutions at the inlet. The swirling flow makes the ball roll round along the edge of the vessel. The velocity of the ball can be measured, for example, by an inductive sensor outside the vessel.After calibration, theflow rate of the fluid can be determined. This concept is called cyclonic flowmeter [9] or orbital ball flowmeter [27] in the literature", + " After determining the stationary solutions of the ball, we identify the parameter values for which the stationary solutions exist (Sect. 3). However, it cannot be done in the case of dual-point rolling due to the undetermined contact forces. This problem can be solved by a new method introduced in Sect. 4, which is applied in the case of dual-point rolling in Sect. 5. Based on these results, the bifurcations of the system are characterised (Sect. 6) and limitations of this type of flowmeters are identified. Let us consider a cylindrical vessel that guides the ball. This is one of the simplest possible geometries (see Fig. 1), which can be found in [9]. Let the ball have a radius of r , and let the inner radius of the cylinder be equal to R + r (see Table 1 for the list of the parameters). 2.1 Kinematic cases Let A denote the downmost point of the ball,where contact can occur with the bottom of the vessel . Similarly, let B denote the outermost point of the ball, which can be in contact with the wall of the vessel. At a contact point between the bodies, three possible situations can occur. The ball can be in rolling or slipping contact with the vessel, or the ball can be separated from the surface of the vessel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003781_j.engfailanal.2020.105095-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003781_j.engfailanal.2020.105095-Figure3-1.png", + "caption": "Fig. 3. Engine exploded view.", + "texts": [ + " The engine damage of 19 users was calculated, and the 95% of the damaged vehicle\u2019s driving route was found. As the target route of acquisition, sensors such as temperature and pressure were installed on the collected field vehicle. The field vehicle drove along the target route and collected the state parameters of engine operation. The linear proportional extrapolation of 2 million km is used as the damage target of engine bench test. According to the principle of equal damage to the user, the engine bench test method is established. Fig. 3 is the engine exploded view. The main reliability durability related failure modes are considered on the market: high cycle Y. Han et al. Engineering Failure Analysis 120 (2021) 105095 fatigue, thermal mechanical fatigue, wear-abrasive, oil deposits, etc. Maximum imposed mechanical load, operating at maximum power and rated speed with a turbo-charged diesel engine, is normally encountered at maximum torque where the cylinder pressure is at a maximum and the lower operating speed, components assessed include: main bearings, piston, liner, crankshaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000692_j.jmatprotec.2015.02.034-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000692_j.jmatprotec.2015.02.034-Figure6-1.png", + "caption": "Fig. 6. The models for the submerged-gate casting process (a) 3D model of the active block product, (b) 3D model of mold cavity that did not need a filling system.", + "texts": [ + " Modeling of casting mold It is necessary to note that the ladle was cuboid with dimensions f 380 mm \u00d7 160 mm \u00d7 160 mm in this application. The relationhip between the volume and the liquid level of the molten metal n the ladle was clear. Thus, there was no need to establish the 3D odel to obtain the slice information. In this section, the 3D model f the mold cavity was required. As mentioned previously, this model successfully avoided a omplicated filling and feeding system. Therefore, the mold design ecame much easier. As shown in Fig. 6, although the casting was omplicated, no filling system was needed in the mold. Only a simle feeder was required to feed the casting and to supply a channel or the sprue to enter into the mold. .2. Slicing of the 3D model By extended development, SolidWorks had the capacity to turn he 3D solid model into slices. In practice, the thickness of the slices ig. 8. The filling process of the submerged-gate casting technology: (a) device and casting etal during the filling process. ig. 9. Casting obtained by the submerged-gate casting technology and its accuracy: (a) c should refer to the height of the whole model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000409_aero.2015.7118995-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000409_aero.2015.7118995-Figure12-1.png", + "caption": "Figure 12. Scheme of sphere in contact with rigid flat before the contact and during the contact.", + "texts": [ + " The expressions for the local stresses and deformation of two spheres in purely elastic contact can be expressed by a Hertz type solution. This solution is equivalent to the solution of a single elastic sphere having an equivalent elastic modulus (E\u2032) and an equivalent radius (R), in contact with a rigid flat surface. The definition of the equivalent parameters are 1 E\u2032 = 1\u2212 \u03bd21 E1 + 1\u2212 \u03bd22 E2 (2) 1 R = 1 R1 + 1 R2 (3) where E1, \u03bd1, R1 and E2, \u03bd2, R2 are the Young\u2019s modulus, Poisson\u2019s ratio and radii of the sphere 1 and 2, respectively. The interference (see Fig. 12), w, is defined as the normal distance that the sphere is displaced into the rigid flat surface. The interference, w, for the elastic case is a function of pmax and defined as w = (\u03c0pmax 2E\u2032 )2 R (4) The yield strength coefficient, C, according to Jackson and Green [11] is defined as C = pc Sy = 1.295 exp(0.736\u03bd) (5) where pc is the critical contact pressure that causes to initial yielding in material and Sy is the yield strength. The Poisson\u2019s ratio, \u03bd, corresponds to the material which yields first" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000657_2013-01-0294-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000657_2013-01-0294-Figure8-1.png", + "caption": "Figure 8. (a) Piezoelectric transducer positioning on engine surface, (b) Instrumented cylinder block.", + "texts": [ + " A series of cut-down 10MHz piezoelectric transducers, having dimensions of 7.1mm \u00d7 1mm were bonded to the exterior surface of the engine, located above the thrust, neutral and anti-thrust sides (15 on each). To enable their mounting, a window was machined through the water jacket at the specified locations and a small flat machined axially along the exterior surface of the cylinder. Each sensor required its own co-axial cable which was sealed around the sensor using a high-temperature, waterproof epoxy and routed out of the engine using specially manufactured covers (Figure 8a and b). An ultrasonic pulser-receiver (UPR) and digitizer unit contained within a PC, was used to pulse and receive the ultrasonic signals (on the same piezoelectric element) and controlling software was written in LabVIEW. The transducers were connected to the UPR via a multiplexer which could programmatically switch between them. The pulse duration applied to the transducers was 100ns (coinciding with their central frequency) at a rate of 80kPulse/s (a rate 20 times greater than that of [25])" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002262_978-3-319-27333-4-Figure3.11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002262_978-3-319-27333-4-Figure3.11-1.png", + "caption": "Fig. 3.11 Lubrication tank with cooling unit", + "texts": [ + " 20/AICTE/ RIFD/RPS(POLICY-II)77/2012-13) and is also compact in design. The new load capacity is 2000 N and speed at which the test rig can run is 6000 rpm. The load arrangement has been changed to pneumatic loading. Further, the test rig is now capable of measuring pressure developed in addition to the measurement of oil temperature rise during the operation. Figure 3.9 shows the new test developed and Fig. 3.10 represents the pressure and temperature sensors. The lubricant supply is made using the lubricant unit as shown in Fig. 3.11. The test rig is equipped with cooling of lubricant during operation time and helps in keeping the lubricant temperature minimal, making it suitable for re-circulation. Example: The performance parameters like pressure, oil-film temperature, Sommerfeld number, power loss and load capacity for an orthogonal journal bearing computed through computer programming, and oil-film pressure variation obtained experimentally have been given below (Figs. 3.12, 3.13, 3.14, 3.15, 3.16, and 3.17). 3.3 Experimental Procedure 32 3 Performance Parameters 35 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002333_978-3-319-41468-3_6-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002333_978-3-319-41468-3_6-Figure12-1.png", + "caption": "Fig. 12 The dimensions of the tested samples", + "texts": [ + " 9 The material ratios of the surfaces (means \u00b1 SD) Fig. 10 Average roughness of the raw samples Fig. 11 The material ratios of the surfaces The material ratios Rmr according to the ISO 13565-2 can be seen in Figs. 9 and 11 (similar results). Machining and mechanical polishing of the samples according to the standards for fatigue testing (Ra 0:8 lm) proved to have a very beneficial effect on the surface quality\u2014Figs. 8, 9, 10 and 11, Tables 6 and 7. A specification of the specimens can be seen in Fig. 12 and Table 4. Fatigue testing was performed using the servo-hydraulic dynamic testing machine Instron 8874\u2014 Fig. 13. Testing frequency of 0.2 Hz in the lower range of testing machine was set because of unknown prediction of cycles to failure. In each series of samples different values of load amplitudes\u20147.7, 6.8, and 6.0 kN with the cycle asymmetry of R \u00bc 1 were applied. The number of cycles to failure N corresponds to a final fracture of the specimens. Results of the series C were rejected because a sheared crack initiated at the contact with gripping jaws that resulted in a premature failure in the followed testing\u2014(see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003324_s11249-020-01316-7-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003324_s11249-020-01316-7-Figure2-1.png", + "caption": "Fig. 2 Schematic of ball-on-disk wear test system", + "texts": [ + " The annealed PV, P0, and P45 samples were marked as APV, AP0, and AP45, respectively. For comparison, a commercial-grade cold-rolled SS316L plate of 3\u00a0mm thickness was also used for the wear test. Specimens of size 20 \u00d7 20 \u00d7 3\u00a0mm were cut from the plate for the wear tests. These specimens were designated as CR. 1 3 The wear resistance of the SLMed and cold-rolled specimens was evaluated using the ball-on-disk testing method (Tribometer, J&L Tech, Republic of Korea). This is shown schematically in Fig.\u00a02. The apparatus comprised a rotating shaft arrangement, to which a disk specimen holder was attached. All the specimens were mechanically ground with SiC papers of up to 2000 grit and swiped with alcohol prior to the wear test. The tests were conducted at 25\u00a0\u00b0C under three normal loads (5, 20, and 50\u00a0N) with a steel ball as a counterpart material. The balls were made of a medium-carbon tool steel with a carbon content of 0.45 wt.% (KS SM45C/AISI 1045). The radius of the ball, Rb, was 5.56\u00a0mm. The maximum Hertzian contact stresses from the above-mentioned experimental set-up were 627, 994, and 1350\u00a0MPa, respectively, for the applied normal loads" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.27-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.27-1.png", + "caption": "Figure 9.27 First kind of motion singularity of a Scara manipulator.", + "texts": [ + "5 as follows: ?\u0307?5 = \u2212\ud835\udf14\u2217 2 (9.328) If s\ud835\udf035 \u2260 0, the same equation gives the following values for ?\u0307?6 and ?\u0307?4. ?\u0307?6 = \u2212\ud835\udf14\u2217 1\u2215s\ud835\udf035 (9.329) ?\u0307?4 = \u2212\ud835\udf14\u2217 3 \u2212 ?\u0307?6c\ud835\udf035 = (\ud835\udf14\u2217 1c\ud835\udf035 \u2212 \ud835\udf14\u2217 3s\ud835\udf035)\u2215s\ud835\udf035 (9.330) A Scara manipulator may have two distinct kinds of motion singularities, which are described and discussed below. (a) First Kind of Motion Singularity Equations (9.323) and (9.324) imply that the first kind of motion singularity occurs if s\ud835\udf032 = 0, i.e. if \ud835\udf032 = 0 or \ud835\udf032 = \ud835\udf0b. This singularity is illustrated in Figure 9.27 in its extended (\ud835\udf032 = 0) and folded (\ud835\udf032 = \ud835\udf0b) versions. Note that the appearance of the folded version of this motion singularity becomes the same as the appearance of the first kind of position singularity if it happens that b2 = b1. In this singularity, i.e. when \ud835\udf0312 = \ud835\udf031 + (1 \u2212 \ud835\udf0e\u20322)\ud835\udf0b\u22152 with \ud835\udf0e\u20322 = sgn(c\ud835\udf032), Eqs. (9.318) and (9.319) can be manipulated into the following forms. (b1?\u0307?1 + \ud835\udf0e\u20322b2?\u0307?12)s\ud835\udf031 = \u2212w1 (9.331) (b1?\u0307?1 + \ud835\udf0e\u20322b2?\u0307?12)c\ud835\udf031 = w2 (9.332) The above equations become consistent and thus they can give finite values for " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002759_2013.17855-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002759_2013.17855-Figure3-1.png", + "caption": "Figure 3. Vehicle Coordinate System", + "texts": [ + " But as the robot tractor moves, it can acquire three-dimensional range information. MAPPING ALGORITHM Coordinates Transformation The data from the LRF are described by polar coordinate system P . Then, it translates those data to vehicle coordinate system which is Cartesian coordinate system. Here, vehicle coordinate system is defined as right-handed coordinate systems of which x-axis correspond to vehicle\u2019s direction, z-axis is downward vertical line, and y-axis is set as to be vertical to the other axes and indicate right direction to the vehicle (Figure 3). Using this translated data, a local map is built which partially represents surrounding scene. The example of local map is shown in Figure 4. [ ]{ }180,1,0, L=\u2208== jPdS T ijiji ijij ss \u03be [ ]{ }180,1,0, L=\u2208== jVzyxL T ijijijij ijij ll = \u03b3\u03be \u03be \u03b3\u03be sinsin cos cossin ijij ijij ijij ij ij ij d d d z y x (1) \u03b3 ; angle between scanning plane and horizontal plane Next we set global coordinate system , according to WGS-84. The transformation from vehicle coordinate to global coordinate is given as follows: G [ ]{ }niVzyxL T iii L1,0, =\u2208== ii ll [ ]{ }niGANEW T iii L,1,0, =\u2208== ii ww ( ) PClTRw ii +\u2212= \u03b8 (2) \u2212 = 100 001 010 T \u2212 = 100 0cossin 0sincos ii ii \u03b8\u03b8 \u03b8\u03b8 \u03b8R [ ]Trrr 321=C [ ]Tiii ZYX=P Where, vector is a position vector of the GPS antenna on vehicle coordinates, while vector P is described by global coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003261_tii.2020.3004389-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003261_tii.2020.3004389-Figure15-1.png", + "caption": "Fig. 15 Simplification and partitioning of the slot winding region.", + "texts": [ + " Below 220\u2103, the variation of thermal conductivities for each material with temperature can be linearized as ( ) 0 0i i i k = + \u2212 (17) where ki is the temperature coefficient derived from the curve fitting of the instrument measurement results at the maximum temperature intended. In general, the stator is the main area for electromagnetic energy conversion. The hotspots are proven to locate in the windings. Unfortunately, the insulation material between wires is vulnerable to the excessive temperature rise [30, 31]. Therefore, it poses a much higher requirement on accuracy [30]. Due to the uneven distribution of the winding ac losses, the slot winding is divided into four subsections and modeled separately to improve accuracy. As shown in Fig. 15, each subsection contains six layers of wires in the block. Fig. 16 presents the equivalent thermal circuit model of a single winding subsection. The end-winding and iron teeth are also divided in the thermal model to establish a complete relationship for thermal transfer. For each component, the variable C and U represent the thermal capacity and heat generation rate respectively, and the variable R represents the thermal resistances between the nodes. Rm, Rn, and Re are the radial resistance between the slot winding layers, the lateral resistance between the slot winding center and the iron tooth, and the axial resistance between the slot winding and endwinding, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure5-1.png", + "caption": "Figure 5. Universal joint.", + "texts": [ + " In Pattern 2, the magnetic pole of the magnet faces the roller side yokes that form a magnetic circuit through the magnetic rollers and adsorption surface. The friction between the roller and the adsorption surface is small and allows the magnet adhesion mechanism to move easily. In addition, because the adhesion yokes are attracted to the magnet by leaked magnetic flux, which is generated by the gap between the adhesion and the rotation yokes, the adhesion mechanism separates from the adsorption surface. The structure of the universal joint attached to each unit is shown in Fig. 5. This joint is used at the connection between each unit and its neighbor. In addition to rotation about the z-axis, controlled by the servomotor, the universal joint gives the units two degrees of freedom. Moreover, the axis of the motor and the two axes of the joint intersect at one point. Each axis of the joint rotates between \u201313\u00b0 and 13\u00b0. Therefore, this robot is flexible on a curved surface. Although the connection is underactuated, the unit moves on a definite trajectory because this robot is constantly on the spherical surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002017_1.4033364-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002017_1.4033364-Figure2-1.png", + "caption": "Fig. 2 Friction measurement test setup", + "texts": [ + " Zhou and Hoeprich extended Aihara\u2019s friction torque calculation by using a different approach to the rib friction calculation and by enhancing raceway roller friction calculation. Qian [14] carried out a study on dynamic simulation of a cylindrical roller bearing. In the study, he used multibodydynamic simulation software to model the bearing completely considering roller pocket contact stiffness, cage geometry, and elasticity of cage. Moreover, validation of analysis is performed with tests. In order to measure the running torque of the slewing ring, a test setup is developed (Fig. 2). The test setup consists of a direct current servo motor, two gearboxes, a coupling, a torque sensor, a bearing unit, and a pinion gear. The motion is transferred to the slewing bearing with the pinion gear supported with a bearing and coupled with the torque sensor. The torque sensor is supported with a leaf spring in order to prevent the rotation of its case. Motion is transferred from servo motor to the torque sensor by means of a coupling. The coupling couples the motion by allowing some angular and radial displacements between the torque sensor and servo motor, caused by manufacturing tolerances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003802_s12541-020-00438-1-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003802_s12541-020-00438-1-Figure9-1.png", + "caption": "Fig. 9 ZMP and foot positions with respect to the GFCF", + "texts": [ + "act y \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (3) Fref z = Fz,L + Fz,R = mg + Kp,height ( zref c \u2212 zact c ) \u2212 Kd,heightz\u0307 act c vertical force of both feet and the reference ZMP derived in Sect.\u00a03.2. First, the x-ZMP, px of the robot can be obtained as in Eq.\u00a0(4). In the equation, xR/L is the right or left foot position in x-direction with respect to the ground-fixed coordinate frame (G.F.C.F.), Fz,R\u2215L is the vertical ground reaction force and MPitch,R\u2215L is the pitch ground reaction moments of the right and left foot, respectively (see Fig.\u00a09). In the same way, the y-ZMP, py can be expressed as in Eq.\u00a0(5), and the total vertical force of both feet can be represented as Eq.\u00a0(6). The ZMP and the total vertical force derived through Eqs. (4\u2013(6) can be expressed as the relationship between the vertical force and the roll/pitch moments of each foot as shown in Eq.\u00a0(7). (4)px = ( xLFz,L +MPitch,L ) + ( xRFz,R +MPitch,R ) Fz,L + Fz,R (5)py = ( yLFz,L +MRoll,L ) + ( yRFz,R +MRoll,R ) Fz,L + Fz,R (6)Fz = Fz,L + Fz,R (7) \u23a1\u23a2\u23a2\u23a3 1 1 0 0 0 0 xL \u2212 px xR \u2212 px 1 1 0 0 yL \u2212 py yR \u2212 py 0 0 1 1 \u23a4\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 Fz,L Fz,R MPitch,L MPitch,R MRoll,L MRoll,R \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 = \u23a1\u23a2\u23a2\u23a3 Fz 0 0 \u23a4\u23a5\u23a5\u23a6 1 3 Equation\u00a0(7) can be briefly expressed as in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000388_romoco.2015.7219734-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000388_romoco.2015.7219734-Figure7-1.png", + "caption": "Figure 7: HoverEye Tail-sitter of Bertin Tech.", + "texts": [ + " Tilt-bodies and Tilt-rotors/Tilt-wings are the two main classes of convertible MAVs. A larger classification of VTOL vehicles, independently of their size, can be found in [1] but as indicated above, some full- scale aircraft have not been built at the MAV scale. Tilt-bodies are essentially airplanes with sufficient thrust to sustain stationary flight. The XFY Pogo with its classical planar wings and SNECMA\u2019s Coleoptere with its annular wing are early examples of this type of aircraft. Many MAVs of this type have been built, like the HoverEye of Bertin Tech. (Fig. 7), the V-Bat of MLB Company (Fig. 8), the MAVion of ISAE (Fig. 9), or the Quadshot of Transition Robotics (Fig. 10). These systems are often referred to as \u201dtailsitters\u201d, due to their capacity to take-off and land on their tail. As indicated by the name, the main characteristics of tilt-bodies is that transition form hover to cruising flight requires the whole body to tilt (pitch down). This implies large variations of the angle of attack, which is one of the control difficulties associated with this type of system. For ducted-fan vehicles like the HoverEye and the V-bat, thrust is generated by one or two (contra-rotative) propellers located inside the duct. Torque control in hover is typically obtained via control surfaces located below these propellers, usually inside the duct so as to preserve flow interactions with the wind (see Fig. 7). Torque control in cruising flight may be aided by additional control surfaces, like for the V-Bat tail-sitter. For the MAVion, the two contra-rotative propellers and control surfaces generate thrust and torque at both hover and cruising flight. In the case of the Quadshot, the four propellers are sufficient to produce both thrust and full torque control. For more efficiency, torque control in cruising flight is also aided by control surfaces on the wing. Besides the problem of transition between hover and cruising flight, a drawback of tail-sitter systems concern payload and aerodynamic stability in hover" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001175_tmag.2014.2356593-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001175_tmag.2014.2356593-Figure3-1.png", + "caption": "Fig. 3. Electric field distribution of helicopter live-line work platform when performing energized work on 1000 kV transmission lines. (a) Helicopter live-line work platform. (b) Line worker.", + "texts": [ + " The potentials applied on the 1000 kV transmission lines in the simulation are prescribed considering the instantaneous value of the three-phase maximum phase voltage, which is 816.5 kV on phase A, and 408.8 kV on phase B and C. The number of degrees of freedom in finite element model is \u223c1.28 \u00d7 107. The calculation is carried out on a workstation with 16 GHz \u00d7 2.40 GHz processors and 64 GB RAM and cost \u223c1 h. The electric field distribution of helicopter live-line work platform when performing energized work on outside phase of 1000 kV transmission lines is shown in Fig. 3. The calculation results indicate that the maximum electric field of helicopter live-line work platform is 44.03 kV/cm, which occurs at the end of helicopter airscrew. The maximum electric field on the surface of line worker is 14.55 kV/cm. The electric field on the head of line worker is within 3 \u223c 5 kV/cm. According to relative standards, the maximum electric field on the body surface of liner worker inside conductive clothing should be no larger than 15 kV/m and the shielding efficiency should be >40 dB for conductive clothing used at nominal voltage up to 800 kV [6]\u2013[8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000469_isie.2015.7281526-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000469_isie.2015.7281526-Figure3-1.png", + "caption": "Fig. 3. Reference frame for the CDFIM with Power Machine stator flux control.", + "texts": [ + " STATOR FLUX ORIENTED CONTROL The torque of CDFIM (equation 8), involving the Power Machine flux and Control Machine stator CUlTent can be written in general reference frame as: (11) This equation shows that the control of torque has a compo nent which is directly proportional to the Power Machine flux and the Control Machine stator CUlTent. By imposing the Power Machine stator flux as the reference frame for the CDFIM vector control, the individual dq-components of the flux space vector in \"e\" reference frame is given by: (12) Where, the superscript \"e\" indicates the excitation ref erence frame of the direct Power Machine flux. Reference frame for the CDFIM in Power Machine stator flux control are depicted in figure 3. Isolating the rotor current in equation 4 and substituting in equation 5 obtains: -e ( LrLsp ) -,e Lr -e -,e 1/Jr = Mp - Mp zsp + Mp 1/Jsp - Mczsc (13) Isolating the Power Machine stator current in equation 5 and replacing it in equation I obtains the model for the CDFIM in stator flux oriented vector control gives in dq components [16]. d{t e Hsp 1/J\ufffdr HspMe i\ufffdse u\ufffdsp - =w =---------- + -- dt K 1/J'dsp K 1/J'dsp 1/J'dsp with K is given by: A. current control loop (15) (16) Rewriting the equation 6 in function of the rotor flux, Power Machine flux and Control Machine CIDTent obtains: (17) with, (18) Replacing the equation 17 in the Control Machine voltage (3) in the vector control reference frame and considering only steady state in dq-components gives: R -e (e ) (Me \"I,e K -e ) seZqse - W - Wepe Ka \ufffddr - bZdse (e ) M pMe \"I,e + W - Wepe K L \ufffddsp a se (19) (20) These equations show two terms of cross coupling, the second term constitute smooth cross coupling in the Control Machine voltage and the third term is a slip proportional term" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003595_icuas48674.2020.9214014-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003595_icuas48674.2020.9214014-Figure3-1.png", + "caption": "Figure 3. Definitions of Coordinates and State Variables", + "texts": [ + " 6-DOF Motion Descriptions Generally, two reference frames are applied to describe the 6-DOF motion of a rigid body. Cartesian coordinates with North-East-Down vector basis { , , }E E EX Y Z denotes the inertial frame fixed on the earth. Cartesian coordinates with Roll-Pitch-Yaw vector basis { , , }b b bX Y Z denotes the body frame attached to the vehicle with its original point at the center of gravity. With the coordinates defined, we can determine the following kinematic variables to describe the UAV\u2019s movement (See Fig.3). [ ]T E E E Ex y z=p denotes the position with respect to the inertial frame, [ ]T E xE yE zEV V V=V denotes the velocity in the inertial frame and [ ]T b xb yb zbV V V=V denotes the velocity in the body frame, [ ]T =\u03be denotes the Euler angles associated with the body frame axis. In this paper, \u2018ZXY\u2019 rotation order is adopted to define Euler angles. [ ]T b p q r=\u03c9 denotes the angular velocity relative to the body frame. Moreover, rotation matrix E bR from body frame to inertial frame and the derivatives of Euler angles are given as: 0 1 / 0 / E b b C C S C C S S S S C S C S C C C S S S C S S S C S C S C C C S S T C T S C C C \u2212 + + = + \u2212 + \u2212 = \u2212 = R \u03be Q\u03c9 Q (1) where C, S, and T respectively denote cosine, sine and tangent functions for short", + " Remark2: To prevent the gimbal lock problem of a tail sitter UAV, \u2018ZXY\u2019 rotation order is used to define Euler angles rather than \u2018ZYX\u2019 which is more common in attitude description. In this paper, we perform the transition along a line trajectory. In this case, the roll angle is around zero, and Q defined by (1) is hence far away from singularity (i.e. 90 = ). [ ( ) ( ) ( )]T wb xb yb zbD t D t D t=D represents the wind speed in the body frame which is generally time-varying and regarded as disturbance to the system. Then, airspeed wV , angle of attack (AOA) \u03b1 and sideslip angle \u03b2 are given by (See Fig.3): ( ) ( ) ( ) ( ) ( ) ( ) 22 2 cos / ,0 tan / , 2 2 w xb xb yb yb zb zb zb zb w yb yb xb xb V V D V D V D V D V V D V D = \u2212 + \u2212 + \u2212 = \u2212 \u2212 = \u2212 \u2212 \u2212 (2) Excluding gravity, resultant force and moment vector acting on the aircraft are denoted as bF and bM in the body frame, respectively. b T m d b fan cs af gyro aero = + + = + + + + F F F F M M M M M M (3) By considering the rotating fan as a steady fan disc acting a thrust and a torque on the aircraft, we can derive the dynamics of the ducted fan rigid body according to Newton and Euler\u2019s Law of Motion: ( )1 1 0 0 E E TE E b b b b b b b g m \u2212 = = + = = p V V R F \u03be Q\u03c9 \u03c9 I M + I\u03c9 \u00d7\u03c9 (4) 1330 Authorized licensed use limited to: Middlesex University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000004_iros40897.2019.8968055-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000004_iros40897.2019.8968055-Figure3-1.png", + "caption": "Fig. 3. The virtual stiffness model based on CP. The direction of hip flexion is defined as the positive direction in this paper.", + "texts": [ + "2(a1) shows swing foot catches up with CP (xCP < xToe) at 69% of swing phase and this moment (tb) is the time of human regaining balance [21]. The distance (\u2206CP ) between CP and swing foot is given by \u2206CP = { xCP \u2212 xToe , if xCP > ToeX xCP \u2212 xHeel , if xCP < HeelX (9) where xToe and xHeel are the x-axis position of toe and heel of swing foot, respectively. ToeX and HeelX are the x-axis position of toe and heel of support foot, respectively. ToeX and HeelX are the upper and lower boundary of BoS, respectively. This paper proposes a virtual stiffness model shown in Fig.3 to generate a virtual torque \u03c4v . And \u03c4v(t) is used as the profile of human-exoskeleton interaction torque \u03c4he(t). The virtual torque \u03c4v acting on the human hip joint is defined as \u03c4v = Kv\u2206CPMh = \u03c4he, (10) where Kv (N/kg) is the virtual stiffness , Mh (kg) is the mass of human. As shown in Fig.3, the virtual stiffness model is a virtual spring connecting CP and the toe or heel of swing foot. When CP exceeds the upper bound of BoS (xCP > ToeX), the virtual spring is connecting to the toe of swing foot, and if CP exceeds the lower bound of BoS (xCP < HeelX), the virtual spring is connecting to the heel of swing foot. The stiffness of virtual spring is denoted as Kv . The virtual spring will generate a virtual torque (\u03c4v) at the hip joint of swing leg. The virtual stiffness Kv is given by (11). When CP is in the BoS there is Kv = 0, and this means virtual spring will not work when the human is in the dynamic balance state. And when CP is in outside of BoS Kv > 0, this means virtual spring is working when human is losing balance. { KV = 0, if HeelX \u2264 xCP \u2264 ToeX KV > 0, if xCP < HeelX||xCP > ToeX (11) As shown in Fig.3, the human body is modeled as an 8187 Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 04,2020 at 14:12:08 UTC from IEEE Xplore. Restrictions apply. inverted pendulum model with finite sized foot [19]. The total mass of the human body is concentrated on the CoM. The position of CoM is approximated at the hip joint of the support leg. The CP is given by xCP = xCoM \u2212 x\u0307CoM/\u03c90, (12) where [xCoM , x\u0307CoM ]T is the state of CoM, \u03c90 is the eigenfrequency of inverted pendulum [19]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001824_1.4033034-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001824_1.4033034-Figure3-1.png", + "caption": "Fig. 3. Variables describing the motion of the wheelset. The sequence of the rotations by the Euler angles \u03c8, \u03d1 and \u03d5 are determined by (7).", + "texts": [ + " Let f\u0302\u00b1w (uw,vw) = u ( f\u00b1w (uw,vw) ) denote the surfaces of the wheels in a general position, where u(r) is the general rigid body transformation which gives the displacement of a point r of the wheelset by u(r) = 1 0 0 0 cos\u03d1 \u2212sin\u03d1 0 sin\u03d1 cos\u03d1 \u00b7 cos\u03c8 \u2212sin\u03c8 0 sin\u03c8 cos\u03c8 0 0 0 1 \u00b7 cos\u03d5 0 sin\u03d5 0 1 0 \u2212sin\u03d5 0 cos\u03d5 \u00b7 r+ x y z . (7) Here, x,y,z denote the displacement of the centre of the wheelset in longitudinal, lateral and vertical directions, respectively, and \u03d1,\u03c8,\u03d5 denote the Euler angles (see Fig. 3 and Tab. 2). According to the convention of railway dynamics, \u03d1 is called roll angle, \u03c8 is the yaw angle and \u03d5 is the rotation angle of the wheelset around its axle. The velocity of any point r of the wheelset can be given by v(r) = u\u0307(r) = \u03d1\u0307\u2212 \u03d5\u0307sin\u03c8 \u03d5\u0307cos\u03c8cos\u03d1\u2212 \u03c8\u0307sin\u03d1 \u03d5\u0307cos\u03c8sin\u03d1+ \u03c8\u0307cos\u03d1 \u00d7 r+ x\u0307 y\u0307 z\u0307 , (8) where the dot denotes differentiation with respect to time. The geometric and kinematic state of the wheelset is restricted by constraints from the rails and from the vehicle. In a general position of the wheelset, let (u+w ,v + w ) and (u\u2212w ,v \u2212 w ) denote the location of the contact points on the surface of the wheels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003261_tii.2020.3004389-Figure22-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003261_tii.2020.3004389-Figure22-1.png", + "caption": "Fig. 22 Installation of monitoring thermocouples in the PMSM.", + "texts": [ + " The three-phase input current is recorded by high-precision current transducers. To obtain the temperature of hotspot in the PMSM and testify the prediction accuracy of the proposed model, two sets of Kthermocouples are embedded at the end and active winding parts of the PMSM. Fig. 21 presents the installation position of the thermocouples in the windings. Besides, to verify the local accuracy of the model, a sufficient number of thermocouples are placed on the iron core, housing, endcap, end cavity, and bearing of the motor, as shown in Fig. 22. Most thermocouples are inserted by drilling small holes and buried by thermal grease. The readings of the thermocouples are recorded by a data acquisition board (Smacq-PS2016). To implement this algorithm for thermal management, the proposed theoretical model is first constructed and debugged in MATLAB. This kind of simulation software can efficiently shorten the development cycle. Before loading into the controller of the thermal management system, the whole model is recompiled to C language and transcoded into the SPEEDGOAT (real-time simulation system) to observe whether it can meet the requirements of the real-time calculation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003073_iet-epa.2019.0941-Figure15-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003073_iet-epa.2019.0941-Figure15-1.png", + "caption": "Fig. 15 Distribution of the local temperature at 12,500 rpm (a) Winding of WT-IPMSM, (b) PM of WT-IPMSM", + "texts": [], + "surrounding_texts": [ + "On the basis of the above analysis, a 45 kW prototype machine with the \u2207 shape PM and water cooling is manufactured and tested. The stator and rotor are shown in Fig. 17 and experimental test platform is shown in Fig. 18. In Fig. 19, the no-load back-EMF at 50 Hz is measured by dragging the proposed WT-IPMSM with two parallel branches. The experimental results are in good agreement with the calculated values. In the load operating characteristic test, the efficiency distributions between the proposed WT-IPMSM and conventional IPMSM are compared in Fig. 20. From Fig. 20a, the measured efficiency values of two prototype motors at 2000 rpm (I) are relatively consistent, which is 91.6%. For Fig. 20b, the number of parallel branches of motors is two when operating at 4000 rpm (II), so they are considered to be the same type of IPMSM, and the efficiencies are 92.7 and 92.6%. At 8000 rpm, the efficiency of the 1192 IET Electr. Power Appl., 2020, Vol. 14 Iss. 7, pp. 1186-1195 \u00a9 The Institution of Engineering and Technology 2020 Authorized licensed use limited to: University of Massachusetts Amherst. Downloaded on July 05,2020 at 09:26:27 UTC from IEEE Xplore. Restrictions apply. proposed WT-IPMSM with four parallel branches (93.2%) is slightly lower than conventional IPMSM (two parallel branches) (94.5%). The lower efficiency for the proposed WT-IPMSM results from larger iron loss. When running at 12,500 rpm (IV), the efficiency of the proposed WT-IPMSM (91.2%) is slightly higher than that of conventional IPMSM (90.6%), due to the fact that the copper loss of the proposed WT-IPMSM is much lower than that of the conventional one. Fig. 21 shows experimental result comparison of torque and current. When the two motors operate at 2000 rpm (I) with the phase current of 310 A, the output peak torque of the proposed WT-IPMSM is close to 3TN (TN = 107.4 Nm) and the conventional IPMSM is only 1.8TN. Therefore, the output torque capacity of the proposed WT-IPMSM is better, which is more suitable for the acceleration and climbing requirements of EV. The proposed IPMSM (a = 2) is the same type of motor as the conventional IPMSM (a = 2) at 4000 rpm (II), with output torque of 107.4 Nm and phase current of 177 A. When the two motors operate at 8000 rpm (III), the output torque of the two motors is 53.7 Nm, and the phase current of the proposed WT-IPMSM is higher than that of the conventional one. The reason is that the number of parallel branches of the proposed WT-IPMSM increases from a = 2 to 4, resulting in a lower back-EMF than conventional one (a = 2). When the two motors operate at 12,500 rpm (IV), the output torques of the proposed WT-IPMSM and the conventional IPMSM are 34.4 Nm, their phase currents are 137.3 and 138.1 A, respectively. When the phase currents of the two prototype motors are the same, the magnetic-field weakening degree of the proposed WT-IPMSM is smaller than that of the conventional IPMSM and its current advance angle is also smaller than that of the conventional IPMSM. Therefore, the q-axis current utilisation rate of the proposed WT-IPMSM is higher, so that it has the ability to further increase the range of speed regulation with field weakening. The compared results between measured and designed values of the proposed WT-IPMSM are given in Table 4. It is obvious that the designed current is slightly smaller and the designed efficiency is slightly higher, but the performance index of suggested design scheme is basically consistent with experimental results. In the load test of the proposed WT-IPMSM, the prime motor applies speed control while the load motor uses torque control. By gradually loading to the designed torque index, load experiments were performed. In addition, the proposed IPMSM and conventional IPMSM are tested in the factory. Owing to the fluctuation of the power grid voltage, the test data includes errors. During design and manufacturing process: (i) the design value of the electromagnetic torque of the proposed WT-IPMSM is slightly smaller. Since the designed Ld and Lq of IPMSM have certain deviations from the real values, in the prototype test, the phase current of the IPMSM was increased to meet the load torque requirements. (ii) Owing to the discrepancy between real value and IET Electr. Power Appl., 2020, Vol. 14 Iss. 7, pp. 1186-1195 \u00a9 The Institution of Engineering and Technology 2020 1193 Authorized licensed use limited to: University of Massachusetts Amherst. Downloaded on July 05,2020 at 09:26:27 UTC from IEEE Xplore. Restrictions apply. predicted value of the friction loss for prototype motor resulting from machining and assembly, the phase current of the IPMSM is increased to meet the load torque requirement. The temperature rises of the 45 kW prototype with the proposed and conventional method are tested to the sub-regional operation under water-cooling condition. Furthermore, the temperature of the winding end is obtained by measuring platinum resistance elements embedded. Clearly, a good agreement can be found between predictions and measurements in Tables 5 and 6. Therefore, the correctness of the electromagnetic\u2013thermal iteration method and thermal simulation model proposed in this paper is validated. The maximum deviation (5.7\u00b0C) between measured and calculated winding temperature occurs in high-speed operation range (IV) of the conventional IPMSM. The reason is that there is an error between the calculated value and actual value of winding copper loss of IPMSM in the case of high-speed operation, and the actual winding copper consumption is higher." + ] + }, + { + "image_filename": "designv11_34_0002302_b978-0-12-397199-9.00004-5-Figure4.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002302_b978-0-12-397199-9.00004-5-Figure4.3-1.png", + "caption": "Figure 4.3: Maximum Range of 4 and z.", + "texts": [ + " Then, using the arc-length formula to compute the maximum length traveled by the agents and invoking its relation with the central angle w, we obtain w\u00f0s\u00de Rtf s k _v\u00f0s\u00dekds r hx \u00fe hy tf s r \u00bc w\u00f0s\u00de; for s\u02db\u00bdtd; tf : Note thatw\u00f0td\u00de \u00bc \u00bd\u00f0hx \u00fe hy\u00ded=r \u00bc qwhilew\u00f0tf\u00de \u00bc 0.We are now left to find4\u00f0s\u00de. To this end, let us analyze when 4 achieves its maximum. We see that the larger admissible value for 4 increases as the error betweenv(s) and vk(s) increases. Accordingly,4(s), s\u02db [tk,tk\u00fe1], can only achieve its maximum when s/t k\u00fe1/tk \u00fe Tk (i.e. right before yk is updated). That is, 4 is upper bounded by the larger possible angle between x(tk) y(tk) and x(tk\u00fel) y(tk). Therefore, let us consider the diagram in Figure 4.3(a), which illustrates not only the angle between x(tk) y(tk) and x(tk\u00fel) y(tk), but also the angle between x(tk) y(tk) and x(tk\u00fel) y(tk\u00fe1). Note that 4 can also be upper bounded by the maximum permissible angle between x(tk) y(tk) and x(tk\u00fel) y(tk\u00fe1), represented by z in the diagram. Hence, let us compute sups\u02db\u00bdtk;tk\u00fe1 kz\u00f0s\u00dek instead. We start by noticing that the difference v(tk\u00fel) v(tk) is upper bounded by kv\u00f0tk\u00fe1\u00de v\u00f0tk\u00dek \u00bc kx\u00f0tk\u00fe1\u00de x\u00f0tk\u00de \u00fe y\u00f0tk\u00fe1\u00de y\u00f0tk\u00dek hxTk \u00fe hyTk; (4.15) which indicates that z is maximized when the equality in (4.15) persists, i.e. when v(tk\u00fe1) lies on the boundary of the ball centered at v(tk) with radius Tk(hx \u00fe hy). Therefore, consider the diagram in Figure 4.3(b), which details this case and chooses e1 and e2 as an orthonormal basis for the plane containing this cross-sectional area of the ball. Let e2 be oriented in the same direction of v(tk). Then, we can define v\u00f0tk\u00de \u00bc kv\u00f0tk\u00deke2. Similarly, v(tk\u00fe1) can be written as v(tk\u00fel)\u00bc c1e1\u00fe c2e2, where c1 and c2 are the new coordinates in the (e1,e2) system. From (4.15), we see that c1 and c2 must satisfy the following equation: c21 \u00fe \u00f0c2 kv\u00f0tk\u00dek\u00de2 hx \u00fe hy 2 T2 k \u00bc r2kv\u00f0tk\u00dek2: Likewise, we observe that z is maximized when the ratio jc1=c2j attains its maximum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003415_s12206-020-0718-y-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003415_s12206-020-0718-y-Figure3-1.png", + "caption": "Fig. 3. Experimental SP equipment and SP test jig setup.", + "texts": [ + " The high temperature SP tests were repeated four times under the same conditions for each block, in this study. SP specimens were machined by wire cutting of each rectangular block. The size of each SP specimen was initially 10\u00d710\u00d7 0.6 mm. The SP specimens were polished on both sides to obtain flat and smooth surfaces by grinding with 800 grade SiC abrasive paper. The final thickness of the specimens was 0.5\u00b10.007 mm. The final geometry of the specimen was 10\u00d7 10\u00d70.5 mm. The jig for the SP test is shown in Fig. 3. The diameter of the loading ball was 2.38 mm. The lower die hole has 4 mm diameter. Edge of the hole was chamfered with corner radius of 0.2 mm. The SP test jig was in compliance with the European Code of Practice [24]. SP testing was conducted at 425 \u00b0C using an INSTRON Universal (E2-016) test machine with a 100 kN capacity. The load was applied to the specimen via the loading ball with a constant loading rate of 0.2 mm/min. During the test, the punch load and displacement data were automatically recorded in the form of plotted curves and displayed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure11-1.png", + "caption": "Fig. 11. Element volume of the SEMMT-applied part on the cutting surface perpendicular to the (a) xy plane, (b) zx plane, and (c) zy plane.", + "texts": [ + " The element volume outside the SEMMT-applied part may be smaller than that inside the SEMMT-applied part. The SEMMT stiffening constant C(e) is chosen such that the elements in the SEMMTapplied part are stiffer than any element outside the SEMMT-applied part. In the present study, the volume of the largest element in the SEMMT-applied part was calculated to be 6.2 \u00d7 10\u22124 and the volume of the smallest element outside the SEMMT-applied part was calculated to be 5.2 \u00d7 10\u22126. Therefore, C(e) is set to 1000 in the SEMMT-applied part, and is set to 1.0 outside the SEMMT-applied part. Fig. 11a, b, and 11c show the distribution of the element volume near the SEMMTapplied part. The red-dashed line in the figures indicates the boundary of the SEMMT-applied part. The outside of the SEMMT-applied part is divided into three bodies, as shown in Fig. 12. In each body, the unstructured elements are generated such that the element volume increases toward the outside to prevent the mesh distortion caused by small elements. This also decreases the number of elements and nodes, decreasing the computational cost" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002100_icciautom.2016.7483208-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002100_icciautom.2016.7483208-Figure1-1.png", + "caption": "Fig. 1. Planar interception geometry", + "texts": [ + " In other words, the main objective of this paper is to propose a guidance law which has an appropriate performance against high maneuvering target motion. Rest of this paper is organized as follows: In section 2, equations of interceptor-target engagement are provided. In section 3, basic definitions of fractional order calculus are given. In section 4, fractional order terminal sliding mode control law will be obtained. Simulation results are shown in Section 5. Finally, conclusion remarks are given in section 6. The geometry of planar interception is shown in Fig. 1. According to the principle of the kinematics, the corresponding equations of motion between the target and the interceptor can be described by [26]: Fractional Order Sliding Mode Guidance Law: Improving Performance and Robustness 2016 4th International Conference on Control, Instrumentation, and Automation (ICCIA) 27-28 January 2016, Qazvin Islamic Azad University, Qazvin, Iran 978-1-4673-8704-0/16/$31.00 \u00a92016 IEEE 469 2 T R MRR R a a\u03bb= + \u2212 (1) 2 T Ma aR R R R \u03bb \u03bb\u03bb \u03bb= \u2212 + \u2212 (2) where R denotes relative distance between the target and the interceptor; \u03bb represents the LOS angular rate; TRa and MRa denote the target and the interceptor acceleration along the LOS, respectively; and Ta \u03bb and Ma \u03bb denote the target and the interceptor acceleration normal to the LOS, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003808_1.g005098-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003808_1.g005098-Figure3-1.png", + "caption": "Fig. 3 Interaction geometry between the platform P and the virtual UAV U.", + "texts": [ + " This procedure can be decoupled by determining the accelerations aUX and aUY of the point U and then using Eq. (3) to determine the inputs to U1 and U2. Thus, the collaborative maneuvering strategy can be developed by interpreting the point U as a single virtual UAV and then designing the control inputs to this virtual UAVas guidance laws, such that the link arrives at the delivery platform with a velocity equal to that of the platform. As the 2-UAV system is now reduced to a single point UAV, to derive the guidance law, the planar engagement geometry, shown in Fig. 3 between the virtual UAV U and a moving platform P, is used, where all angles are measured positive counterclockwise about the positive X axis. The virtual UAV can maneuver in the 2-D space by changing the magnitude aU and direction \u03b3U of its acceleration. These controls result in a change in its flying speed VU and heading angle \u03b1U. Similarly, the maneuvering platform can also change its speed VP > 0 and heading angle \u03b1P, by changing the magnitude aP and direction \u03b3P of its acceleration. Defining the separation distance between U and P as r, the LOS angle as \u03b8, and following the kinematic relations shown in Fig. 3, the nonlinear kinematics-based interaction equations are as follows: _r Vr VP cos \u03b1P \u2212 \u03b8 \u2212 VU cos \u03b1U \u2212 \u03b8 (4a) r_\u03b8 V\u03b8 VP sin \u03b1P \u2212 \u03b8 \u2212 VU sin \u03b1U \u2212 \u03b8 (4b) _VU aU cos \u03b3U \u2212 \u03b1U (4c) VU _\u03b1U aU sin \u03b3U \u2212 \u03b1U (4d) _VP aP cos \u03b3P \u2212 \u03b1P (4e) VP _\u03b1P aP sin \u03b3P \u2212 \u03b1P (4f) These equations are subsequently used to derive the sliding-modebasedmaneuvering strategy or guidance law, in terms of aU and \u03b3U as controls. These controls appear in Eqs. (4c) and (4d) in a nonaffine form. To enforce sliding mode, it is preferable that these controls appear in an affine form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002875_s42496-020-00033-7-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002875_s42496-020-00033-7-Figure3-1.png", + "caption": "Fig. 3 MSC Adams solar arrays equivalent model", + "texts": [ + " The aforementioned equations have been used to obtain the sloshing parameters to be implemented in a commercial suite to simulate the dynamics of the system. The system has been modelled using a multibody dynamics software, namely MSC Adams. This is a Multibody Dynamics (MBD) commercial tool used to analyse articulated mechanical systems based on the principles of Lagrangian Dynamics. The code can solve both algebraic and differential non-linear equations as functions of time. As far as the solar arrays are concerned, a realistic system would also include the presence of hinges, torsional springs and dampers to connect consecutive sections (see Fig.\u00a03). Due to the high computational costs related to such model in the Adams\u2013Simulink architecture and the necessity to perform several dynamic simulations, an equivalent monolithic structure has been adopted to carry out the analyses. According to such an approach, the allowed relative rigid rotations between each section of the panels are reproduced as an increased flexibility of the overall array. (4)mi = mf [ tanh ( 2 ih / ) i ( 2 i \u2212 1 ) h ] , (5)hi = h 2 \u2212 2 i [ tanh ( ih ) \u2212 1 \u2212 cosh ( 2 ih / ) sinh ( 2 ih\u2215 ) ] , (6)ki = mig 2 i tanh ( 2 ih ) , (7) 1 = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000584_j.jsc.2014.09.031-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000584_j.jsc.2014.09.031-Figure7-1.png", + "caption": "Fig. 7. The other case of fold by (O5).", + "texts": [ + " We will postpone the description of our proof until Section 5 since the proof scheme is the same in other constructions. We move on to the construction of a pentagonal knot. On the tape shown in Fig. 4, we perform a fold (O5) along a fold line that passes through F, to superpose H and CG. We have two fold lines, say m1 and m2 to make this possible. The fold along m1 creates the shape as shown in Fig. 6. Point K is created at the intersection of m1 and line CG (in Fig. 4). The other case of the fold (along m2) constructs the shape shown in Fig. 7(b), where point K is not constructible. We now see how to fold the desired knot. The following shows the Eos program to construct the pentagon EHGKF. Program P1 (Construction of a pentagonal knot). 1. BeginOrigami(\u201cPentagonal knot\u201d, {150, 10}); 2. NewPoint({E \u2192 {45, 0}, F \u2192 {43, 10}}); 3. HO(\u2203m,m:Line\u2203n,n:Line\u2203g,g:Point\u2203h,h:Point\u2203i,i:Point(O5(g, EA, E, m) \u2227 m = EF \u2227 O5(F, EB, g, n) \u2227 n = gh \u2227 g \u2208 CD \u2227 h \u2208 AB \u2227 i \u2208 DmF \u2227 i \u2208 Cn g), MarkPointAt \u2192 {G, H, I}, MarkPointOn \u2192 {{CG, K}}, Handle \u2192 {A, B}); 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000822_wcica.2014.7053425-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000822_wcica.2014.7053425-Figure6-1.png", + "caption": "Fig. 6 Block diagram of the proposed controller in Z-domain So, the differential equations of the controller can be written as:", + "texts": [ + " The step response and response to a disturbance of each controller in simulation in s-domain are shown in Fig.5, where 3pK , 1iK , 0.003D , 2J , pK . The value of the disturbance is 0.2 N m . It is easy to find that when the value of becomes higher, the transient response of the controller becomes faster, but the peak response becomes higher as well. While the settling times and the responses to the disturbance of the controllers are the same. The block diagram of the proposed controller in z-domain is indicated in Fig.6. ( )R z represents the reference input, ( )V z stands for the feedback signal, while ( )I z is the output of the controller, and ( )U z is the intermediate variable. pK is the proportional coefficient, iK represents the integral coefficient, is the feedforward gain. 1z stands for the pure delay element. Using the anti-windup strategy in [14], the integral state of the controller would not windup when the reference changes, a better performance can be got. The block diagram of P-I or I-P controller can be got by making the value of equals 0 or pK " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure6-1.png", + "caption": "Figure 6. Contraction of the cam mechanism.", + "texts": [ + " In addition to rotation about the z-axis, controlled by the servomotor, the universal joint gives the units two degrees of freedom. Moreover, the axis of the motor and the two axes of the joint intersect at one point. Each axis of the joint rotates between \u201313\u00b0 and 13\u00b0. Therefore, this robot is flexible on a curved surface. Although the connection is underactuated, the unit moves on a definite trajectory because this robot is constantly on the spherical surface. The contraction of the cam mechanism is shown in Fig. 6. First, the magnetic pole faces the adhesion yokes, and the unit is extended and fixed on the surface (Fig. 6 (a)). The unit stays expanded as the cam mechanism moves to Fig. 6 (b), while the magnet rotates through 90\u00b0 to face the roller side yokes. In this configuration, the friction plates rise, and the unit can move. Next, as the configuration changes from Fig. 6 (b) to Fig. 6 (c), the unit is contracted without rotation of the magnet. Thus, the unit is easily contracted. Extension of the cam mechanism is performed in the reverse order of contraction. The modeled shape of the robot for the controlling omnidirectional locomotion is shown in Fig. 7. The contraction links and joints are conjoined in an octagonal configuration. In the previous research [12], we modeled the shape of the robot on flat ground, because we expected the robot to move only on flat surfaces. In this research, we model the robot in three-dimensional (3D) space, in the light of motion on spherical surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000233_978-1-4939-7247-0-Figure6.3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000233_978-1-4939-7247-0-Figure6.3-1.png", + "caption": "Fig. 6.3 Design of an ankle-foot orthosis stiffness testing device. The novel Biarticular Reciprocating Universal Compliance Estimator (BRUCE) design provides an excellent example of a clinically viable device that provides valuable quantifications of orthosis bending stiffness. (Figure originally published in Bregman et al. [51])", + "texts": [ + " Notably, one group has developed a novel, clinically applicable device to test the stiffness and neutral angle of AFOs [51]. The Biarticular Reciprocating Universal Compliance Estimator (BRUCE) measures stiffness of an AFO and/or AFO-shoe combination about the ankle and metatarsal joints across a functional sagittal plane ankle range of motion. It accommodates a wide range of AFO sizes and shapes because it measures angles and moments in the same manner as a traditional gait analysis and applies boundary conditions that mimic conditions during AFO use. BRUCE can also be operated in a clinical setting (Fig. 6.3). Other techniques have been tested to measure AFO stiffness, but many do not apply appropriate boundary conditions. This is potentially problematic because AFOs are very sensitive to loading conditions [52]. In addition to stiffness, other important characteristics have been identified as necessary for customized orthoses. For example, footplate design and orthosis joint alignment are two components that can be customized to improve gait function by facilitating foot-to-floor motion and controlling kinematics of the entire lower extremity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001163_detc2013-12748-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001163_detc2013-12748-Figure3-1.png", + "caption": "Fig. 3 A scheme of experimental apparatus", + "texts": [ + "65 JIS G4051 S45C Thermal refining steel Non-profile-shifted 4 5 29 / 29 Hobbing 116 10 4 20 116 3 Copyright \u00a9 2013 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/77583/ on 03/04/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 1x105, 5x105, 1x106, 2x106, 5x106, 1x107 2x107, and 3x107 in order to observe the tooth surface. Then, the changes of the tooth profile, the tooth surface roughness and wear loss are measured. Figure 3 shows the power circulating-type gear testing machine of which total length is about 2 meters. The test gear box and bearing stand are separated not to be impaired by the wear powder generated from test gear unit and suporing bearings. The test gears are specified in Table 1 and shown in Fig. 4. The test gears have 4 mm module, a standard 20\u00b0 involute pressure angle, 29 teeth, and 10mm face width. They are made of thermal refining steel JIS G4051 S45C. The Vickers Hardness of test gears is about 220 Hv" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001175_tmag.2014.2356593-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001175_tmag.2014.2356593-Figure6-1.png", + "caption": "Fig. 6. Experimental arrangement of helicopter live-line work by platform method on 1000 kV transmission lines.", + "texts": [ + " An experiment of helicopter live-line work by platform method on 1000 kV transmission lines is performed to verify the safety of line worker and the validity of simulation results. The 1000 kV transmission lines are replaced by a conductor with the same structure, and the helicopter liveline work platform is placed at a supporting platform, which is 2.5 m \u00d7 2.5 m \u00d7 0.15 m in size and 12 m in height. The supporting platform is supported by four poster insulators, and the experiment system well imitates the floating situation of helicopter. The illustrations of the arrangement of experiment are shown in Fig. 6. During the experiment, the electric potential of conductor is increased to 635 kV (the effective phase voltage of 1000 kV transmission line in consideration of over-voltage) gradually. The equipotent bonding of helicopter live-line work platform with conductor is realized by extending the bonding wand. The arc discharge of conductor to bonding wand is observed during the experiment, as shown in the elliptic region in Fig. 6. The experimental results indicate that the equipotent bonding procedure has no influence on helicopter and line worker. The electric field right, the head of line worker was measured by the electric field probe for three times to improve the accuracy of the measurement results, and the average value obtained is 2.06 kV/cm. The environment temperature during the experiment is 23.2 \u00b0C, and the environment humidity is 51.3%. The discharge current during the bonding procedure has been measured, and the sampling frequency is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001773_s00419-016-1123-y-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001773_s00419-016-1123-y-Figure2-1.png", + "caption": "Fig. 2 Amplitude\u2013frequency dependence of the softening Duffing oscillator according to perturbation analysis or harmonic balance (Eqs. 4, 5), exact solution in case of \u03b7 = 0 (blue dots) for parameters D = 0.06, \u03b5 = \u22120.1 and f = 0.2. (Color figure online)", + "texts": [ + " Please notice that in most argumentations afterward, this relation (3) will not hold! As a standard method to obtain the solution of (2) harmonic balance is used. In the simplest case the ansatz x(\u03c4 ) = C cos (\u03b7\u03c4 \u2212 \u03d5) (4) is performed. Inserting this into (2) and neglecting the higher-order frequency term proportional to cos 3\u03b7\u03c4 , the amplitude C and phase shift \u03d5 can be obtained by solving the nonlinear algebraic equations ( 1 \u2212 \u03b72 + 3 4 \u03b5C2 )2 C2 + 4D2\u03b72C2 = f 2, tan \u03d5 = 2D\u03b7 1 \u2212 \u03b72 + 3 4\u03b5C 2 . (5) These results for the amplitude C are shown in Fig. 2. Beside the well-known behavior in resonance with the curve having an inclination to the left and the occurrence of multiple solutions (denoted as solutions 2\u00a9), additional \u201dnose-like\u201d solutions for small \u03b7 can be found with large amplitudes C (denoted as solutions 1\u00a9). For weaker (but non-vanishing) damping and/or larger excitation amplitude, the \u201dnose-like\u201d solutions and the inclined resonance peak converge and are finally merging. Now, the oncoming considerations will focus on the accurateness of these \u201dnose-like\u201d solutions and\u2014after having found several arguments for the nonaccurateness\u2014on alternative solutions for the low frequency range", + " Looking at the assumptions made in the beginning, it is stated in Eq. (3) that 1 |\u03b5|x2 should hold. For the \u201dnose-like\u201d solution, this assumption is obviously violated, as we are in the range of x , where the restoring characteristic in Fig. 1 is crossing zero, i.e., 1 \u2248 |\u03b5|x2. Another assumption due to Eq. (7) was that \u03b7 \u2248 1. Here we have \u03b7 close to zero. Additionally, for \u03b7 = 0 (static case), an exact solution of (2) can be calculated by solving the equation C + \u03b5C3 = f. (8) This result is marked by blue dots in Fig. 2. Obviously the two upper solutions are unstable. There are some deviations between the amplitudes calculated from this and from the perturbation analysis. Using numerical integration, the \u201dnose-like\u201d solutions could not be found as shown in Fig. 3. Of course, this would also happen, if both \u201dnose-like\u201d solutions are unstable. Now, these \u201dnose-like\u201d solutions will be investigated more in detail by adding an error estimation to the harmonic balance method. We are now again considering the harmonic balance method for the solution of (2)", + " In the following, we will refine harmonic balance in that way that we use the neglected terms (11) as a measure for the approximation error. These terms can be calculated from the found solution ak, bk . As an error we choose e\u0302 = max 0\u2264\u03c4\u2264 2\u03c0 \u03b7(n+1) { 3n\u2211 k=n+1 ( a\u0303k cos(k\u03b7\u03c4) + b\u0303k sin(k\u03b7\u03c4) )} (12) which gives the maximum of the neglected term (11) over one period in time. Restricting the order to n = 1 as in the section before, the resulting error e\u0302 is shown in Fig. 4a. Herein the large errors for small \u03b7 belong to the \u201dnose-like\u201d solution (denoted as solution 1\u00a9 in Fig. 2 and afterward). Compared to this, the error of solution 2\u00a9 in Fig. 2 is very small. Enlarging the order to n = 3 in the harmonic balance, additional solutions occur with large amplitudes for small \u03b7 while the behavior for \u03b7 \u2248 1 remains qualitatively unchanged. Those new solutions contain also large errors that are not plotted here. The solutions 1\u00a9 and 2\u00a9 in Fig. 2 change slightly for n = 3, and the corresponding errors are shown in Fig. 4b. It is clearly visible that the solutions with large amplitudes at small \u03b7 1\u00a9 still have large errors while the errors at \u03b7 \u2248 1 for solutions 2\u00a9 have decreased. From these results it can be concluded that the \u201dnose-like\u201d solutions 1\u00a9 seem to be very inaccurate. But how do potential additional solutions in this frequency range look like? We consider again Eq. (1). In contrast to the transformation used to get Eq. (2), we now use the circular excitation frequency \u03a9 for performing the transformation \u03c4 = \u03a9t" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000898_10402004.2013.830799-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000898_10402004.2013.830799-Figure1-1.png", + "caption": "Fig. 1\u2014Radially loaded roller bearing.", + "texts": [ + " It is assumed that the bearings are subjected to pure radial load and there is no roller misalignment. Rollers are supposed to be equispaced and the cage-to-roller interaction force is neglected. The outer race is considered to be fixed to the casing and the inner race rotates with the shaft. The races are assumed to be flexurally rigid and undergo only local deformation. Centrifugal force and gyroscopic moments on the bearing are ignored. The load distribution in a statically loaded roller bearing is determined according to the procedure outlined in Sarangi, et al. (23), (24). Figure 1 shows the radially loaded roller bearing. When the load is applied on the bearing, the rollers occupied in the region 2\u03b81 undergo elastic deformation and with dynamic conditions, the external load will be supported by the lubricant film. The total interference at an angle \u03b8 with respect to the load line is given by \u03b4\u03b8 = ( \u03b40 + Pd 2 ) cos \u03b8 \u2212 pd 2 , [1] where \u03b40 is the total interference along the line of action of a load Wy. Considering the Hertz elastic deformation given by Harris (21) and the Dowson and Higginson film thickness relation (Dowson and Higginson (25)), the expression for \u03b4\u03b8 can be written as \u03b4\u03b8 = KHCF1/1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000370_icma.2015.7237694-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000370_icma.2015.7237694-Figure1-1.png", + "caption": "Fig. 1 The multi-point interaction between the redundant manipulator and environment", + "texts": [ + " (2) The force controller is applied to accomplish the contact task in force control stage. (3) Collision avoidance in both position and force control stage. (4) The trajectory in joint space is closed in condition that the trajectory in Cartesian space is closed So the contact impedance control strategy is utilized in main task compliant surface, while the non-contact impedance control strategy is used in the obstacle avoidance compliant surface (second compliant surface). The dual surface interaction between the redundant manipulator and environment is shown as Fig.1. The dual compliant surface 1424978-1-4799-7098-8/15/$31.00 \u00a92015 IEEE consists of main task compliant surface and obstacle avoidance compliant surface. B The Impedance Character in Main Task Compliant Surface The impedance controller makes the manipulator act as a mass-spring-damping system. And the hybrid impedance control can ensure compliance in position control subspace and accuracy force control in force control subspace. In the main task compliant surface, the objective of hybrid impedance control is as follows ( ) ( ) ( ) ( ) d d d d d d e e e e e e e e e e e e d e e e - + - + - - - =- M X S X B X S X K S X X I S F F (1) with known notation as in [10]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001917_ccdc.2014.6852195-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001917_ccdc.2014.6852195-Figure1-1.png", + "caption": "Fig 1. Morphing aircraft\u2019s simplified model.", + "texts": [ + " The controller\u2019s structure is designed by the combination of the linear quadratic (LQ) output feedback approach to ensure stability under the fixed-wing state and the tracking error PI compensator to keep the steady-state errors zero while tracking the command. The controller\u2019s parameters are optimized by the SRAD method. Simulation results for a morphing UAV demonstrate the effectiveness of the controller. 2 AIRCRAFT\u2019S MATHEMATICAL MODEL 2.1 Physical Model This paper focuses on the morphing aircraft which can change its wings\u2019 span and sweep angle in flight. In the modeling process, the aircraft is simplified into a multi-rigid-body system which is composed of five separate rigid bodies (shown in Fig 1), and their masses are mb, m1, m2, m1 and m2. The total mass of the aircraft is mt. The fuselage is simplified into a cylinder with radius R, length l0, center of mass Cb. The inner and outer parts of the aircraft\u2019s wings are all simplified into homogeneous rods. The six-degree-of-freedom dynamic model is built based on Kane\u2019s Method in reference [13]. Except for the special maneuvering flight, the aircraft\u2019s structure is symmetric about the central plane. So, the aircraft\u2019s dynamic equations can be simplified to a set of simple dynamic equations by setting 1 2, 1 2", + " Consequently, longitudinal dynamic equations are simplified into: 2 2 2 1 1 1 1 1 1 1 1 1 1 1 1 2 2 1 1 1 1 2 1 1 1 1 1 1 1 1 1 2 0 cos sin( ) cos sin cos ( cos sin sin ) sin (2cos sin ) 2 cos ( sin sin sin 2cos cos ) 2 sin (2 t t b b T D m g m V m bq m bq m l q m l q q m L q L L m \u03b1 \u03b8 \u03b1 \u03b1 \u03b1 \u03b1 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b1 = \u2212 \u2212 \u2212 \u2212 + + \u2212 \u2212 \u22c5 + \u22c5 + \u22c5 \u2212 \u22c5 + \u22c5 \u2212 \u22c5 \u2212 \u22c5 \u0394 + \u22c5 \u2212 \u22c5 \u0394 \u2212 \u22c5 \u2212 1 1 1 1 1 1 1sin sin 2 cos )q L q L q\u03b8 \u03b8 \u03b8 \u03b8\u0394 + + (1) 2 2 2 1 1 1 1 1 1 1 1 1 1 1 1 2 2 1 1 1 1 2 1 1 1 1 1 1 1 1 1 2 0 sin cos( ) ( ) sin cos sin ( cos sin sin ) cos (2cos sin ) 2 sin ( sin sin sin 2cos cos ) 2 c t t b b T L m g m V q m bq m bq m l q m l q q m L q L L m \u03b1 \u03b8 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b8 \u03b8 \u03b1 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = + \u2212 \u2212 \u2212 \u2212 + \u2212 \u2212 \u2212 \u22c5 + \u22c5 + \u22c5 + \u22c5 + \u22c5 \u2212 \u22c5 \u2212 \u22c5\u0394 + \u22c5 \u2212 \u22c5 \u0394 \u2212 \u22c5 + 1 1 1 1 1 1 1os (2sin sin 2 cos )q L q L q\u03b1 \u03b8 \u03b8 \u03b8 \u03b8\u0394 + + (2) 1 1 1 2 1 1 1 1 1 1 1 1 1 1 2 1 1 1 1 1 1 0 ( sin cos ) ( sin 2 sin ) cos ( sin + cos cos ) sin (2 sin + cos 2 cos 2 2 cos sin ) 2 sin ( sin + cos cos 2sin sin y b b b D L M bm m l m L g m b V V bq qV lm V V qV l q l q m L V V qV q L \u03b1 \u03b1 \u03b8 \u03b8 \u03b8 \u03b1 \u03b1 \u03b1 \u03b1 \u03b8 \u03b1 \u03b1 \u03b1 \u03b1 \u03b8 \u03b8 \u03b8 \u03b8 \u03b1 \u03b1 \u03b1 \u03b1 \u03b8 \u03b8 = + + + \u2212 + + + \u22c5 \u2212 \u2212 \u2212 \u22c5 \u2212 + \u22c5 + \u22c5 \u2212 \u22c5 \u2212 + \u22c5 \u0394 + 1 1 1 2 2 3 1 1 1 1 2 cos ) 2( )(sin 2sin cos )y q L q J q J J q q \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 + \u22c5 \u2212 \u2212 + \u22c5 + \u22c5 (3) where T is the engine\u2019s trust, D is drag force, L is lift force, My is pitching moment, is pitch angle, q is pitch angular velocity, is the angle of attack, V is the aircraft\u2019s translational velocity, 2 1 1 1 2 l L l= + \u0394 \u2212 , other parameters are shown in Fig. 1. Kinematic equations are simplified into: q\u03b8 = (4) sin( )h V \u03b8 \u03b1= \u2212 (5) where h is the aircraft\u2019s flight altitude. Furthermore, the elevator\u2019s dynamics are modeled as a first-order lag model: e e c K T T \u03b4 \u03b4 \u03b4 \u03b4\u03b4 \u03b4= \u2212 + (6) where e is the elevator\u2019s deflection angle, c is the commanded elevator input. Equations (1~6) are the longitudinal dynamics of the morphing aircraft. 474 2014 26th Chinese Control and Decision Conference (CCDC) 2.3 Aerodynamic Forces The accurate modeling and numerical calculation for morphing aircraft\u2019s aerodynamic coefficients are very complex, and they are also the developing technologies and of the most interesting research areas in morphing technologies [20, 21]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000757_icaccct.2014.7019468-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000757_icaccct.2014.7019468-Figure1-1.png", + "caption": "Fig. 1 Net torque acting on horizontal and vertical plane", + "texts": [], + "surrounding_texts": [ + "ISBN No. 978-1-4799-3914-5/14/$31.00 \u00a92014 IEEE 381\nmultivariable self tuning PID (Proportional-IntegralDerivative) controller (STC) for pitch angle and yaw angle control of the main and tail rotor of the Twin Rotor MIMO System (TRMS) under parametric uncertainty and external disturbance. The PID parameters are updated in real time based on relationship between PID and minimum variance control (MVC). The TRMS model is identified based on an on-line auto regressive moving average (ARMAX) model whose parameters are tuned by a recursive least square (RLS) algorithm. The results obtained show superiority of the proposed controller with respect to a fixed gain PID controller under external disturbances. Keywords\u2014 Self tuning PID, TRMS, MVC, ARMAX, RLS\nI. INTRODUCTION\nHelicopters have been widely used in air traffic systems, especially in urgent transportation needs such as medical Treatment rescues and life-saving goods transportations [1]. In Addition, missions like ground detection, traffic condition Assessment, smuggling prevention, and crime precautions, Fire detection depend on helicopters [2]. A twin rotor MIMO system (TRMS) has been considered as a challenging research topic as it resembles the behavior of a helicopter in certain aspects. With complexity such as nonlinearity and high crosscoupling effect between two rotors. As TRMS is subjected to coupling effects due gives rise to infinite modes of transversal vibration in the beam. A mathematical model under such external forces acting simultaneously is not sufficient to represent the dynamics of the TRMS. Hence there have been numerous attempt in the past to identify the complex dynamics of the TRMS using a discrete time model [3-4]. Also from [5] we know that an indirect adaptive control can be designed using an identified model of any complex system. Also minimum variance based self tuning control uses variance of the tracking error to design a control input has shown good results in controlling a uncertain system [6]; with this motivation in this paper. In this paper a mathematical modelling of the TRMS and then two multivariable adaptive controllers were designed based on on-line identified ARMAX model. The results obtained show superiority of the proposed controller with respect to a fixed gain PID controller and minimum variance self-tuning control (STC) under external disturbances [7].\nII. DYNAMIC MODELLING OF A TRMS\nTRMS is a laboratory set-up used for control purpose. The working principle is based on the aerodynamic laws like helicopter. TRMS consists of two rotors main rotor rotates along vertical plane and tail rotor rotates along horizontal plane. Both the rotors are driven by two D.C. motors. TRMS is a two degree of freedom but the vertical degree of freedom can be restricted to one degree of freedom by nylon screw found near pivot point.\nModelling of TRMS can be derived from the electro mechanical diagram as shown in fig 2. For Main Rotor:\nGravitational Torque: It is produced by the model weight. It acts vertically downward.\n\u03c8\u03c9\u03c4 singrw ML == (1)\nwhere,\nw\u03c4 = Gravitational torque (N-m)\nL = Moment arm (m) \u03c9 = weight of the helicopter (kg) \u03c8 = Elevation angle\ngrM = Gravity Momentum parameter (N-m)\nCentrifugal Torque: It is produced by Centrifugal force during rotation in horizontal plane. It acts vertically downward.\n\u03c8\u03c4 coscc LF= (2)\nwhere,\n\u03c8 \u03c6\nsin 2\n2\ndt\nd mLFc = (3)\ncF = Centrifugal force (N)\n\u03c8 = elevation angle (rad)\ndt\nd\u03c6 = angular velocity in horizontal plane (rad/sec)\nThe main rotor force is a consequence of its angular speed. If angular speed is more than more force will be induced on the system body.\nAngular torque = ( )111 wF\u03b1\u03c4 (4)\nwhere,", + "1F = Main rotor force (N)\n1w = Main rotor angular velocity (rad /sec)\n1\u03c4 = Main rotor torque (N-m) The main motor is approximated by the first order transfer function which can be described as\n1\n1011\n1 1 u\nTsT\nk G\n+ = (5)\nWhere\n1u = Input to the motor1\n1k = constant\nThe nonlinearity caused by the rotor i.e. assumed as a second order polynomial & the torque induced in helicopter body through motor is given as below.\n11\n2\n111 GbGa +=\u03c4 (6)\n11 ,ba = Nonlinearity constant parameters\nGyroscopic Torque: This torque results when main rotor changes its position in azimuth direction. Thus the Gyroscopic Torque\n\u03c8\u03c4 \u03c6\n\u03c4 cos1 dt\nd k gyG = (7)\nWhere\n\u03c6 = Azimuth angle\ndt\nd\u03c6 = Angular velocity in Azimuth\ngyk =Constant of proportionality (sec/rad)\nFrictional torque: It is produced due to air resistance. For main rotor it acts vertically against the air resistance..It is denoted by.\ndt\nd Bf \u03c8 \u03c4 11 = (8)\nWhere\n1B = Damping constant (NMS/rad)\nThe net torque obtained from the Newton\u2019s 3rd law. Thus the net torque act on helicopter is\n112\n2\n1 fwGc dt\nd I \u03c4\u03c4\u03c4\u03c4\u03c4 \u03c8 \u2212\u2212++=\n(9)\nwhere,\n1I = moment of inertia of helicopter along horizontal axis\nG\u03c4 = Gyroscopic torque\nc\u03c4 = Centrifugal torque\nw\u03c4 = Gravitational torque\n1f\n\u03c4 = Frictional torque\nSimilarly the Azimuth dynamics for Tail rotor: The net torque act on helicopter is\n222\n2\n2 fr dt\nd I \u03c4\u03c4\u03c4 \u03c6 \u2212\u2212= (10)\nwhere\n2I = Moment of inertia in vertical plane (kg.m^2) \u03c6 = Azimuth angle\n2\n2\ndt\nd \u03c6 = Azimuthally acceleration\n2\u03c4 = Side rotor torque\n2f \u03c4 = Frictional torque\nr\u03c4 = Main rotor reaction torque\nFrictional Torque: It is denoted by 2f\n\u03c4 .\ndt\nd Bf \u03c6 \u03c4 22 =\n(11)\nwhere,\n2B = Damping constant (NMS /rad)\nMain rotor reaction torque: It is produced due to the coupling between main rotor and tail rotor. It acts along azimuthal direction. It can be estimated by the first order transfer function.\n( ) 1 10 + + = sT sTk\np\nc r\u03c4 (12)\nWhere,\npo TT , = Cross reaction momentum parameter\nck = Constant\nNow the tail rotor is approximated by the first order transfer function which can be described as\n2\n2021\n2 2 u\nTsT\nk G\n+ = (13)", + "Where\n2u = Input to the motor (2)\n2k = Constant\nThe nonlinearity caused by the rotor can be estimated as second order polynomial. The net torque induced in helicopter\nthrough the motor is\n22\n2\n222 GbGa +=\u03c4 (14)\nWhere\n22 ,ba = nonlinearity constant parameters\nEquation (10) and (11) together represent the system dynamics.\nIII. PROPOSED MULTIVARIABLE ADAPTIVE CONTROL DESIGN\nA. Selection of model structure\nA recursive parameter estimator is the most important part\nof the self tuning algorithm. The main idea of recursive\nestimator is to update the system parameters (A, B, C) of the\nmodel structure (ARMAX) continuously [6, 10, 15]. For\nsimplification ARMAX model is chosen.\n( ) ( ) ( ) ( )\n( ) ( ) ( )\n( ) ( )cn\nbn\nan\nntectec\nntubtubtub\nntyatyatyaty\nc\nb\na\n\u2212+++\n\u2212\u2212++\u2212+\u2212+\n\u2212\u2212\u2212\u2212\u2212\u2212=\n,...,\n1,.....,21\n,....21\n0\n10\n21\n(14)\nThe above equation is represented as\n( ) ( )\u03b8\u03c6 1\u2212= tty T (15)\nWhere \u03c6 is called regressor which contains all the past values\nof output y and input and also includes delayed noise terms.\n\u03b8 Includes all the parameters of input, output and noise term which to be estimated. RLS is used to update the parameters of ARMAX model [6, 8, 10].The plant model can be written as\n( ) ( ) ( ) ( ) ( ) ( )dke zcdkuzBdkyzA + +\u2212+\u2206=+\u2206 \u2212\u2212\u2212 111 1 (16)\n( ) ( ) ( ) ( ) T\np tutututu ],.....,,[ 21=\n( ) ( ) ( ) ( ) T\np tytytyty ],.....,,[ 21=\n\u2206 Is the difference operator defined as 11 \u2212\u2212=\u2206 z . (17)\n( ) ( ) ( ) ( )],.....,,[ 11\n2\n1\n1 1 \u2212\u2212\u2212\u2212 = zAzAzAdiagzA p\nIn ARMAX model we consider up to second order polynomial\ni.e. ( ) ( ) ( )2\n2,\n1\n1,\n1 1 1 \u2212\u2212\u2212 ++= zazazA ii . (18)\n( )1\u2212zB is the polynomial matrix represented as\n( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( ) ( )\n( ) ( )m\nm\npppp\np\npiii\nzBzBB\nzBzBzB\nzBzBzB\nzBzBzB\nzB\n\u2212\u2212\n\u2212\u2212\u2212\n\u2212\u2212\u2212\n\u2212\u2212\u2212\n\u2212\n+++=\n \n\n \n\n=\n......1\n10\n1\n,\n1\n2,\n1\n1,\n1\n,2\n1\n2,2\n1\n1,2\n1\n,\n1\n2,\n1\n1,\n1\n(19)\nB. PID controller design\nThe PID controller is designed for the system model for d=3.The PID controller is represented as digitized form as\n( )kei is the control error signal represented as\n( ) ( ) ( )kykrke iii \u2212= ,where r is the desired signal and y is the actual output signal.\n( ) 211 2 11 \u2212\u2212\u2212 +\n +\u2212 ++= z T T kz T T k T T T T kzL s D p s D p s D i s pi\n(21)\nThis can be written as\n( ) ( ) ( ) ( ) ( ) 011 =\u2212\u2206+ \u2212\u2212 trzLtutyzL iiiii (22) Now the above PID control laws is compared with minimum variance control law to derive the relationship between unknown parameters of PID and known parameters of minimum variance control law [15-17].\nThe parameters of minimum variance controller are obtained by solving the Diophantine equation. C. Self tuning controller Fig 3 represents structure of self-tuning PID using MVC. The Diophantine equation is represented as\nGZAFC d\u2212+= (23)\nWhere\n( ) ( ) ( ) 1\n1\n1\n10\n1\n1\n1\n1\n1\n1\n......\n......1\n+\u2212\n\u2212\n\u2212\u2212\n+\u2212\n\u2212\n\u2212\u2212\n+++=\n+++=\nn\nn\nd\nd\nZgZggZG\nZfZfZF\n(24)" + ] + }, + { + "image_filename": "designv11_34_0001507_ijca.2015.8.2.16-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001507_ijca.2015.8.2.16-Figure1-1.png", + "caption": "Figure 1. Joint with Planetary Gear Reducer", + "texts": [ + " A planetary gear dynamic model is consist of the sun gear (subscripts s ), the planet gear (subscripts pi , 1,2i n ), the ring gear and carrier (subscripts c ), where each stage has n planet gears as shown in Figure 2. Copyright \u24d2 2015 SERSC 155 The nonlinear model is established by the lumped parameter method in the following hypotheses: (1) All the planetary gears are assumed to be spur gears with the same physical and geometrical parameters. (2) The influence of friction in the process of the gear mesh can be neglected. Each corresponding mathematical model (Figure 1) is obtained by Newton-Euler method as follows: The analysis of the sun gear: 1 [ ] n s s D spi spi s i J T P D r (1) The analysis of the ith planet gear: +pi pi spi spi rpi rpi piJ P D P D r (2) The analysis of the carrier: 2 1 1 ( ) + + cos( ) n n c pi c c spi spi rpi rpi c L i i J m r P D P D r T (3) Where, J (subscripts s , pi , c ) is the inertia ranged from sun gear to planet carrier, sr is the base circle radius, cr is the carrier radius which is equivalent to the pitch circle radius of sun gear add to the pitch circle radius of planet, m (subscript pi ) is mass of planet gear, is the mesh angle of gears, DT is the input matrix, LT is the load matrix, and P , D (subscripts spi , rpi ) are elastic mesh force and viscous mesh force, in which subscript spi represents the internal meshing between sun gear and the planet gear and subscript rpi represents the external meshing between the planet gear and carrier, respectively", + " Through the above processing, the solving difficulty of nonlinear equations is reduced, at the same time the precise integration algorithm [15] can be used to handle with first-order differential equation. Due to the simple principle of precise integration algorithm, the steady-state response results can be obtained by means of few codes with high run speed. In this section, the differential equations will be solved numerically by precise integration algorithm (integration step is 1ms). The main parameters of the planetary gear transmission joint shown in Figure 1 (the planetary has 3 planet gears) are given in Table 1. In this study, the value of average mesh stiffness mK is 810 , and the initial phase angle of frequency error is assumed to be 0. The phase diagram can be used to analyze the dynamic characteristics due to the core is to analyze the influence of backlash to gear periodic solutions. Assuming that the backlash values are both 2b between the meshing of sun-planets and planet-ring gear, the steady-state response can be calculated under different backlash values, which is characterized using the dimensionless displacement X " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003808_1.g005098-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003808_1.g005098-Figure1-1.png", + "caption": "Fig. 1 Apair ofUAVs connectedusing a rigid link interacting, in the 2-D plane, with a moving platform.", + "texts": [ + " Next, having located the object, simple waypoint-based motions are executed to grasp, transport, and release the object at the designated area. A similar approach is followed by Nieuwenhuisen et al. [19] in locating, grasping, and delivering objects using UAVs. The collaborativemanipulation scenario considered in this paper is where a pair of UAVs connected to each other with a rigid link work cooperatively to retrieve a payload from, or deliver it to, a stationary or a moving platform. A schematic of this application is shown in Fig. 1, where the platform is moving at a constant heading angle and speed, VP. Thus, for this application, for the 2-UAV system to track the platform, the twoUAVshave to ensure that they too fly at the same heading angle and speed as the platform. The use of a rigid link may seem restrictive, but such a system performs the following functions: 1) It removes the need to adopt a collision-avoidance strategy between the UAVs themselves, thus reducing the communication load between them. 2) It allows the considerations of the forces, arising from the interaction between the 2 UAVs, only at the ends of the rigid link, which in turn can be compensated by each UAV", + " In this paper, themaneuvering strategies for the 2-UAVsystem are developed by enforcing sliding mode, using the lateral and longitudinal accelerations of the UAVs as controls, on manifolds defined in the relative velocity space. The manifolds are selected according to the nonsingular terminal sliding mode approach. The motivation to choose nonsingular terminal sliding mode stems from the inherent property of sliding mode occurring on the chosen manifold within a finite-time interval, and so the states that define these manifolds also have finite-time convergent properties. Thus, for the 2-UAV system interacting with the moving platform, as shown in Fig. 1, the sliding manifolds, and corresponding controllers, are designed such that the 2-UAV system can track the platform within a finite-time interval. As will be discussed in the paper, for the case of stationary platforms, with the enforcement of sliding mode on the selected manifolds, the approach angle to such platforms by the 2-UAV system can be specified. Similarly, for the case of tracking a maneuvering platform, with a minor change in the design of the manifolds, the 2-UAV system can track the platform within a finite-time interval" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001730_978-3-319-18296-4_13-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001730_978-3-319-18296-4_13-Figure1-1.png", + "caption": "Fig. 1 Wheel/rail geometry: a reference position; b particular position", + "texts": [ + "), Soft Computing Applications, Advances in Intelligent Systems and Computing 356, DOI 10.1007/978-3-319-18296-4_13 155 This paper focuses on a new geometric model of the contact between a wheel with irregular contour and a smooth rail. This model is more realistic since it takes the wheel curvature into account. The interaction between a flat wheel and a rigid smooth rail is being studied, by using the new geometric model. This chapter deals with a wheel with irregular contour that is rolling without slipping along a perfectly smooth rigid rail, as seen in Fig. 1. The irregular contour of the wheel is described by the function q \u00bc q\u00f0h\u00de; \u00f01\u00de where \u03c1 and \u03b8 are the polar coordinates in respect to the wheel center Ow. Point M on the wheel contour has the coordinates x and y in respect to the orthogonal reference Oxy, where the origin point O belongs to the wheel contour x \u00bc q sin h \u00f02\u00de y \u00bc qo q cos h; \u00f03\u00de where \u03c1o is the distance from the wheel center Ow to the point O. The orthogonal reference Oxy is attached by the head of the rigid rail. The slope of the tangent M\u03c4 is given as tanu \u00bc dy dx : \u00f04\u00de By differentiating relations (2) and (3), it is obtained dx \u00bc \u00f0q cos h\u00fe q0 sin h\u00dedh \u00f05\u00de dy \u00bc \u00f0q sin h q0 cos h\u00dedh; \u00f06\u00de where q0 \u00bc dq dh : \u00f07\u00de Inserting the above differentials in (4), it reads dy dx \u00bc tanu \u00bc q sin h q0 cos h q cos h\u00fe q0 sin h : \u00f08\u00de The geometrical considerations drive to the angle between the OwM and the normal M\u03bd b \u00bc h u: \u00f09\u00de When the wheel rolls without slipping on the rail, the initial contact point O between the wheel and the rail arrives in O\u2032, the point M comes in the point M\u2032 on the rail head; when the wheel center Ow arrives in the point Ow\u2032 (Fig. 1b), the abscissa of the M\u2032 point is xr. We are interested in calculating the coordinate (xw, yw) of the wheel center when the wheel turns around the center point at the angle \u03b8. To this end, the length of the arch OM has to be known. According to the geometrical correlation in Fig. 2, the length of an infinitesimal element of the wheel contour is ds \u00bc q cos b dh \u00f010\u00de and the length of the wheel contour between O and M points is s \u00bc Z ds \u00bc Zh 0 q cos b dh \u00f011\u00de When the wheel rolls on a rigid rail and the wheel rotation angle is \u03b8, the distance between the initial contact point O and the current wheel\u2013rail contact point M\u2032 and the arch length are equal xr \u00bc s: \u00f012\u00de According to Fig. 1b, the coordinates of the wheel center are as follows xw \u00bc xr q sin b \u00f013\u00de yw \u00bc q cos b: \u00f014\u00de In the preceding geometrical considerations, we tacitly suppose that the wheel and the rail have a single contact point. However, this hypothesis has to be verified. To this purpose, the curvature of the wheel contour must be calculated C \u00bc d2 y dx2 1\u00fe dy dx 2 3=2 : \u00f015\u00de The derivation dy/dx is given by Eq. (8), while the derivation of second order can be computed according to the equation d2y dx2 \u00bc d dh dy dx 1 dx dh \u00bc q2 qq0 \u00fe 2q02 q cos h\u00fe q0 sin h\u00f0 \u00de3 : \u00f016\u00de Upon inserting Eqs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002010_lascas.2016.7451026-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002010_lascas.2016.7451026-Figure1-1.png", + "caption": "Fig. 1. Model for the experimental 3D-printed generator. One of the lower combs is also shown.", + "texts": [ + " MEMS implementations of similar generators, powered by mechanical vibrations, were recently demonstrated [4]. In this work, an investigation is made about the construction of variable-capacitance generators using plastics, exploring 3Dprinting techniques. The idea is to obtain medium-sized devices for experimental purposes, which may be also eventually realized in the small dimensions of MEMS devices with high-precision printing technologies that are in rapid development. For the investigations, an experimental device was constructed, with original plans shown in Fig. 1. It comprised two moveable combs and four fixed combs made with conductive ABS plastic, mounted in a structure of normal insulating ABS making a base, two supporting springs, and an attachment for a vibrating device. Metal screws and nuts were used to join the parts and make contacts for the electrical connections. The complete structure could be vibrated, causing the combs to make two complementary variable capacitors, but for a more precise evaluation just the moveable part was vibrated, by a simple shaking mechanism moved by an electric motor", + " The capacitors were dimensioned to have a maximum capacitance of approximately 150 pF, ignoring edge effects, with 15 plates in the lower combs and 14 in the upper combs, all with 13\u00d754 mm, with 1 mm of thickness and of separation, dimensions restricted by the precision of the printing technology used. In the next sections, several possible structures for a generator are discussed, and their performances evaluated. II. GENERATORS BASED ON BENNET\u2019S DOUBLER The electronic version of Bennet\u2019s doubler shown in Fig. 1a, proposed in [2], uses two complementary variable capacitors Ca and Cb, a fixed storage capacitor C1, and three diodes. With the This work was partially supported by the CNPq. ISBN 978-1-4673-7835-2/16/$31.00\u00a92016 IEEE 127 IEEE Catalog Number CFP16LAS-ART experimental generator, it is obtained by connecting the two sections in parallel. This has the same effect of a single complementary variable capacitor with doubled capacitance, with combs two times thicker, but requires just about one-half of the displacement for the same capacitance variation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003753_jestpe.2020.3037942-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003753_jestpe.2020.3037942-Figure1-1.png", + "caption": "Fig. 1. SFPM Machine topologies. (a) Conventional 12s13r SFPM. (b) C-core 6s13r SFPM. (c) E-core 6s13r SFPM. (d) 6s13r CP-SFPM.", + "texts": [ + " In Section III, a hybrid finite element (FE)/analytical modeling is presented to quantify the torque contributions of the multiple air-gap field working harmonics by employing a unified torque equation. In Section IV, the torque contributions of the main-order working harmonics in the four SFPM machines are identified and quantified. Some experimental measurements on a CP-SFPM prototype are carried out in Section V. Some conclusions are drawn in Section VI. A conventional 12-stator-slot/13-rotor-tooth (12s13r) SFPM machine is illustrated in Fig. 1(a). In order to reduce the PM consumption, two 6s13r SFPM machines with C-core and E-core stators [5]-[12] are developed as shown in Figs. 1(b) and (c), respectively. Besides, in order to provide an effective magnetic path for the lower-order working harmonics, a CP-SFPM machine [24] [25] with iron bridge design in stator yoke is designed and depicted in Fig. 1(d). In order to perform a fair comparison, the four SFPM machines share the same stator outer/inner diameters, air-gap length, active stack length, rotor structure and current density, as listed in Table I. For a doubly salient structure, the air-gap permeance can be obtained by employing a slotless rotor and a slotless stator, respectively, which can be expressed as [34] 0 1 1 1g s r g (1) where g is the air-gap length, \u03bc0 is the vacuum permeability. \u039bs and \u039br are the air-gap permeance functions with the slotless rotor and slotless stator conditions, respectively, which can be referred as stator and rotor permeances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001035_09205071.2014.963698-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001035_09205071.2014.963698-Figure1-1.png", + "caption": "Figure 1. Side view of the circular coils of rectangular cross section and parallel axes shielded by two parallel screens of high permeability.", + "texts": [ + " For the coaxial circular coils shielded by the ferromagnetic screens, the mutual inductance is given in [7]. However, as yet the results of the non-coaxial coils shielded by the screens have seemingly not been found in the literatures, therefore it is necessary to discuss this problem in detail. To highlight the most important factors of this problem, the radial dimension and the permeability \u03bc of both screens are supposed to be infinite, and the axes of the coils are set vertically to the surface of the screens. The whole configuration is shown in Figure 1. In this boundary value problem, the Neumann formula of mutual inductance is unusable and the magnetic energy must be considered, so the expressions of the magnetic field of the shielded coils are essential. We will see that the series-integral expansion of the reciprocal distance 1/R in the cylindrical coordinate system is useful for obtaining the proper form of the source magnetic field, which is fundamental for the *Email: ostpreussen@qq.com \u00a9 2014 Taylor & Francis Journal of Electromagnetic Waves and Applications, 2014 Vol", + " r1 2lr2 (21) Consequently we obtain 2 Z1 0 f2 t\u00f0 \u00deJ1 r2t\u00f0 \u00dedt \u00bc r1 2lr2 \u00fe Z W f2 t\u00f0 \u00deH 1\u00f0 \u00de 1 r2t\u00f0 \u00dedt (22) With (8), (19), (22), eventually we obtain M \u00bc l0p l r21 2 \u00fe r1r2 P1 k\u00bc1 cos ak f2 f1\u00f0 \u00de \u00fe 1\u00f0 \u00dek cos ak f2 \u00fe f1\u00f0 \u00de h i I1 akr1\u00f0 \u00deK1 akr2\u00f0 \u00deI0 akq0\u00f0 \u00de ; q0 r2 r1 (23) When 0\\r2 r1\\q0\\r1 \u00fe r2, the analytical continuation cannot be implemented and the mutual inductance must be evaluated with (8) numerically. The mutual inductance of the shielded thick circular coils can be obtained from (8), (18), and (23). Consider two thick coils of parallel axes, one has inner and outer radii R1, R2, axial length 2h1, and N1 turns, another has corresponding parameters of R3, R4, 2h2, and N2. The distance between the axes of both coils is \u03c10, and the middle planes of them are set in z = z1 and z = z2, respectively (See Figure 1). In the case of no overlap of radial arguments between two coils, (18) and (23) can be employed. By performing the integration with respect to the axial arguments of (18) and (23), we have Zz1\u00feh1 z1 h1 Zz2\u00feh2 z2 h2 cos ak f2 f1\u00f0 \u00de \u00fe 1\u00f0 \u00dek cos ak f2 \u00fe f1\u00f0 \u00dedf2df1 \u00bc 4 cos ak z2 z1\u00f0 \u00de \u00fe 1\u00f0 \u00dek cos ak z2 \u00fe z1\u00f0 \u00de h i sin h1ak\u00f0 \u00de sin h2ak\u00f0 \u00de . a2k (24) Carrying out the remaining radial integration of (18) gives M \u00bc l0p 3N1N2 cl P1 k\u00bc1 cos ak z2 z1\u00f0 \u00de \u00fe 1\u00f0 \u00dek cos ak z2 \u00fe z1\u00f0 \u00de h i sin h1ak\u00f0 \u00de sin h2ak\u00f0 \u00de u ak;R2;R1\u00f0 \u00deu ak;R4;R3\u00f0 \u00deK0 q0ak\u00f0 \u00de a4k; q0 R2 \u00fe R4 (25) where c \u00bc 4h1h2 R2 R1\u00f0 \u00de R4 R3\u00f0 \u00de, and u ak; q2; q1\u00f0 \u00de \u00bc q2 I1 akq2\u00f0 \u00deL0 akq2\u00f0 \u00de I0 akq2\u00f0 \u00deL1 akq2\u00f0 \u00de\u00bd q1 I1 akq1\u00f0 \u00deL0 akq1\u00f0 \u00de I0 akq1\u00f0 \u00deL1 akq1\u00f0 \u00de\u00bd (26) Similar integration of (23) gives M \u00bc l0pN1N2 l R2 2\u00feR2R1\u00feR2 1 6 \u00fe p2 c P1 k\u00bc1 cos ak z2 z1\u00f0 \u00de \u00fe 1\u00f0 \u00dek cos ak z2 \u00fe z1\u00f0 \u00de h i sin h1ak\u00f0 \u00de sin h2ak\u00f0 \u00deu ak;R2;R1\u00f0 \u00dev ak;R4;R3\u00f0 \u00deI0 akq0\u00f0 \u00de a4k ; q0 6 R3 R2 (27) Where v ak; q2; q1\u00f0 \u00de \u00bc q2 K1 akq2\u00f0 \u00deL0 akq2\u00f0 \u00de \u00fe I0 akq2\u00f0 \u00deL1 akq2\u00f0 \u00de\u00bd q1 K1 akq1\u00f0 \u00deL0 akq1\u00f0 \u00de \u00fe I0 akq1\u00f0 \u00deL1 akq1\u00f0 \u00de\u00bd (28) For 0\\R3 R2\\q0\\R2 \u00fe R4, the series representations are invalid and we must start with (8)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000004_iros40897.2019.8968055-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000004_iros40897.2019.8968055-Figure1-1.png", + "caption": "Fig. 1. The human-exoskeleton interaction model of single joint. \u03b8h and \u03b8e are the joint angle trajectory of human and exoskeleton, respectively. \u03b8\u0307h and \u03b8\u0307e are the joint velocity trajectory of human and exoskeleton, respectively. \u03c4h and \u03c4e are the joint muscle torque of human and the joint driving torque of exoskeleton, respectively. \u03c4he is the human-exoskeleton interaction torque. lh and le are the length of human limb and exoskeleton, respectively. mh and me are the mass of human limb and exoskeleton, respectively. khe and che are the stiffness coefficient and damping coefficient of humanexoskeleton interaction model, respectively. The anti-clockwise direction is defined as the positive direction.", + "texts": [ + " Based on the concept of integral admittance, the quantitative definitions of assistance and resistance are that when Xhe > Xh the human joint is assisted by the exoskeleton and when Xhe < Xh the human joint is resisted by the exoskeleton. And it is noticeable that if Xhe = Xh human joint is neither assisted nor resisted by the exoskeleton, and this situation is usually called \u2018transparent\u2019. Xh = \u03b8h \u03c4h (1) The key to change the integral admittance of human limb joint is to control the interaction torque (\u03c4he) between human limb and exoskeleton. In order to reveal the primary relationship between the human-exoskeleton interaction torque and the integral admittance, one joint of human limb shown in Fig.1 is analyzed, and the interaction effect between human and exoskeleton is represented by a spring-damper model. Therefore the dynamic equations of the human muscle torque (\u03c4h), the exoskeleton driving torque (\u03c4e) and the humanexoskeleton interaction torque (\u03c4he) are given by \u03c4h = mhglhsin(\u03b8h) +mhl 2 h\u03b8\u0308h \u2212 \u03c4he, (2) \u03c4e = meglesin(\u03b8e) +mel 2 e \u03b8\u0308e + \u03c4he, (3) \u03c4he = khe(\u03b8h \u2212 \u03b8e) + che(\u03b8\u0307h \u2212 \u03b8\u0307e). (4) The control objective of active assistance is to enlarge the integral admittance of human limb joint (Xhe > Xh)", + " In the first way of enlarging the integral admittance of human limb joint (\u03980 h = \u0398\u2032h), according to (2), the human joint torque is given by \u03c4 \u2032h = mhglhsin(\u03b8\u2032h) +mhl 2 h\u03b8\u0308 \u2032 h \u2212 \u03c4he = mhglhsin(\u03b80 h) +mhl 2 h\u03b8\u0308 0 h \u2212 \u03c4he = \u03c40 h \u2212 \u03c4he. (5) Therefore, the human-exoskeleton interaction torque in this situation is given by \u03c4he = \u03c40 h \u2212 \u03c4 \u2032h. (6) The integral admittance of human limb joint when exoskeleton assisting is Xhe = \u03b80 h/\u03c4 \u2032 h, and the integral admittance of human limb joint before wearing exoskeleton is X0 h = \u03b80 h/\u03c4 0 h . If Xhe > Xh and \u03980 h = \u0398\u2032h, it requires \u03c4 \u2032h < \u03c40 h , therefore \u03c4he > 0 which means the direction of human-exoskeleton interaction torque is in the same direction of human joint torque, as shown in Fig.1. Similarly, in the second way of enlarging the integral admittance of human limb joint (\u03c4 \u2032h = \u03c40 h). According to (2), the human joint torque when exoskeleton works in active assistance mode is given by \u03c4 \u2032h = \u03c40 h = mhglhsin(\u03b8\u2032h) +mhl 2 h\u03b8\u0308 \u2032 h \u2212 \u03c4he = mhglhsin(\u03b80 h) +mhl 2 h\u03b8\u0308 0 h. (7) Therefore, the human-exoskeleton interaction torque in this situation is given by \u03c4he = mhglh(sin(\u03b8\u2032h)\u2212 sin(\u03b80 h)) +mhl 2 h(\u03b8\u0308\u2032h \u2212 \u03b8\u03080 h). (8) If Xhe > Xh and \u03c4 \u2032h = \u03c40 h , it requires \u0398\u2032h > \u03980 h, and according to equation (8) there is \u03c4he > 0 which also means the direction of human-exoskeleton interaction torque is in the same direction of human joint torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003932_s40962-020-00549-5-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003932_s40962-020-00549-5-Figure1-1.png", + "caption": "Figure 1. Types of turbo-expander wheel: (a) semi-open type, (b) enclosed type, (c) open type.", + "texts": [ + ", a constant-entropy process), and the low-pressure exhaust gas from the turbine is at an exceptionally low temperature. It accounts for 80*90% in the entire process of producing the cold temperatures. The turbo-expander wheel of the large oxygen plant must have high dimensional accuracy and good dynamic stability. Because it transfers external work to the rotating shaft, it also must have enough strength and plasticity for operating at room or cryogenic temperature. In general, there are three types of turbo-expanders for oxygen plants; semi-open, enclosed and open type (Figure 1). The advantage of a semi-open type and an open type is that their manufacturing process is comparatively simple and are light. However, their efficiency is lower than that of an enclosed type design, which is higher, but its manufacturing process is rather complicated.1 However, there is strong demand to increase the efficiency and quality of the wheel and reduce the time and cost of its manufacture by looking at new production approaches. Additive manufacturing (AM) technology, often referred to 3D printing or rapid prototyping (RP), directly fabricates the complex geometrical part (concept models, functional prototypes and final products) by layer-by-layer deposition of the materials from CAD 3D model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002708_3033288.3033299-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002708_3033288.3033299-Figure1-1.png", + "caption": "Figure 1. The boat model", + "texts": [], + "surrounding_texts": [ + "Plant identification is done by adopting the nonlinear autoregressive exogenous model (NARX) approximation, expressed as: (1) where y is the plant output, u is the plant input, and are the delay or memory operators for the plant output and input, respectively. In this work, the neural networks based system identification is a multi-layer perceptron neural network which consist of one input layer with 15 neurons, one hidden layer with 30 neurons, and one ouput layer with 3 neurons, respectively. The configuration for the boat model system identification is shown in Figure 5. All of the neurons other than that in the input layers use a bipolar Sigmoid activation function. To obtain the plant model, a back-propagation learning mechanism is adopted to train the neural network configurations with learning rate equal to 0.2. After 701,595 epochs, the training converge with a training mean-sum-square error (MSSE) of 2.2383 x 10-4. The testing results of the system identification are depicted in Figure 6. The testing MSSE is 3.679 x 10-4, which can be detailed as mean-square-error (MSE) of each output parameters as follows: the MSE for yaw is 1.811 x 10-4, the MSE for vx is 1.838 x 10-4, and the MSE for vy is 7.388 x 10-4. These low values of errors reflect that the designed neural networksbased identification system can successfully model the real transfer function of the plant with very high approximation." + ] + }, + { + "image_filename": "designv11_34_0002056_1350650116631453-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002056_1350650116631453-Figure2-1.png", + "caption": "Figure 2. Forces acting on the roller in the loaded zone.", + "texts": [ + " According to conformability of deformation, the deformation can be described as25 \u00bc max cos \u00f02\u00de In accordance with the relationship between the deformation and the contact load, one equation can be attained as25 Q Q0 \u00bc max 1=t \u00f03\u00de where t is the coefficient, t\u00bc 3/2 for ball bearings, t\u00bc 10/9 for roller bearings. Equations (2) and (3) are substituted into equation (1) which can be expressed as follows Q0 \u00bc W ZJ J \u00bc 1\u00fe2 P cost Z ( \u00f04\u00de where 1=J is a constant correlating with the roller numbers Z (see Houpert27 and Wan25). The loading conditions in the loaded zone and unloaded zone are given in Figure 2 and 3, respectively. Qij is the normal force between the roller and inner race, Qoj represents the normal force between the roller and outer race. Taking 1=J \u00bc 4:6, the normal force between the maximum loaded roller and inner race is given by27,25 Qi0 \u00bc 4:6W Z \u00f05\u00de The normal contact force between the maximum loaded roller and outer race is described as Qo0 \u00bc Qi0 \u00fe F! \u00f06\u00de where F! is the centrifugal force of the roller, which is written as F! \u00bc mRm! 2 c \u00f07\u00de where m is the mass of roller. When bearing is operating under lubrication condition, hydrodynamic pressure cannot be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000584_j.jsc.2014.09.031-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000584_j.jsc.2014.09.031-Figure10-1.png", + "caption": "Fig. 10. Intermediary steps of heptagonal knot construction.", + "texts": [ + " We can start the construction of a heptagonal knot by the application of Isosceles Trapezoid Lemma (Lemma 2). Then, we pick the point B to pull the faces on which B is on, and perform a mountain fold (O5), then pick A and perform a mountain fold (O5), and finally pick A and perform a valley fold (O5). The parameters of the folds (O5) are chosen to construct the congruent isosceles trapezoids. We need to pay attention to the keyword arguments Direction (Mountain or Valley) and InsertFace (above Bottom or below Top), to make the knot rigid. Fig. 10 shows the intermediary steps of the fold of heptagonal knot. We finally obtain the heptagon ELHGJKF shown in Fig. 11. Similarly to the case of the regular pentagonal knot, to make the polygon ELHGJKF a regular heptagon, we need to add further constraints of the equalities of the edges. We redo the construction after adding the constraints. With hands, we pull the edges AD and BC outwards, as we did with the regular pentagonal knot. This will move point F and perturb the whole shape. The final one is shown in Fig. 12. As for the proof for the construction shown in Fig. 10, we give to the prover, before specifying the goal, the following assumptions: SquaredDistance(E,F) = SquaredDistance(G,H) = SquaredDistance(F,K) = SquaredDistance(E, L) = SquaredDistance(J,G); (10) SquaredDistance(E,F) = SquaredDistance(K, J) = SquaredDistance(L,H). (11) The assumption (10) comes from Isosceles Trapezoid Lemma and the assumption (11) is the one ensured by the fastening of the tape. By both assumptions we intend to assume the equalities among the length of the edges involved. The assumption (10) should be unnecessary theoretically, since this is the consequence of Isosceles Trapezoid Lemma" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002574_detc2016-60019-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002574_detc2016-60019-Figure6-1.png", + "caption": "Fig. 6 Illustration of contour generation procedures (a) Slicing list (b) After sorting neighbor list (c) after arrangement of the head and tail status of list (d) final contour list", + "texts": [ + " When all the possibilities of intersection between the slicing plane and the corresponding facets are checked, and corresponding intersection points are computed, we must connect the points in each facet to generate the straight-line. This slicing list array stores the lines which intersect with the slicing plane in a random order. Based on the slicing list array the closed contour is generated by the contour generation algorithm. Firstly, the algorithm sorting the first segment from the slicing list array to a new file then finds its neighbor or a follow line. When the required follow line has discovered by the program, it will be recorded in a new file, but it might be possible to get an unordered list as shown in Fig. 6. For these straight-line data, there will be two points. The start point is called as head of straight-line and end point is termed as the tail of the straight-line. In general, the fully closed contour is generated only if the head of one straight-line lies on the tail of neighboring straight-line. According to head-to-tail search as shown in Fig. 6, we can find out the correct arrangement of head and tail status of neighbor straight-line to connect the straight-lines. For each slice height, this procedure is repeated until all lines are connected. Using this comparison process for the whole slicing list, we can generate the fully closed contour [30-34]. 4 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/90683/ on 02/05/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use For the Data transfer in coupled operations, MATLAB 14 programming language is used as a source system to write the contour lists file" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002410_s12649-016-9708-9-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002410_s12649-016-9708-9-Figure5-1.png", + "caption": "Fig. 5 The dependence of the concentration of WSN and NH4\u2013N in water on the particle size of HM in the composite films used for the solubility tests. Particle size of HM: 0.40\u20130.25 mm (1), 0.25\u20130.16 mm (2), 0.16\u20130.09 mm (3), B 0.09 mm (4)", + "texts": [ + " In addition to organic nitrogen horn meal contains small amount of water soluble nitrogen (WSN), which makes ca. 0.5 % of the total amount of nitrogen [42, 43]. The decrease of particle size of horn meal leads to the increase of the surface of their contact with water, which results in the increase of the solubility. The experiments of the estimation of the amount of WSN in water, which was used for the evaluation of the solubility of the composite film showed that the decrease of the particle size of the filler leads to the increase of the amount of WSN (Fig. 5). The amounts of WSN in the films of the samples 1, 2, 3 and 4 was 0.035, 0.057, 0.087 and 0.111 % respectively. The amount of ammonium nitrogen (NH4\u2013N) in WSN was 23.3, 26.6, 21.0 and 19 % respectively. The increase of the amount of the soluble compounds in the solutions leads to the increase of the pH values from 6.2 to 6.6 (Fig. 6). Water vapour permeability of the composite films was estimated. It was found to be 0.618 9 10-12, 0.446 9 10-12, 0.383 9 10-12 and 0.306 9 10-12 mol/m 9 s 9 Pa for the films of the composites 1, 2, 3 and 4 respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001841_pcitc.2015.7438180-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001841_pcitc.2015.7438180-Figure1-1.png", + "caption": "Figure 1. Schematic model of three-phase SEIG for WECS", + "texts": [ + " For controlling purpose, the study of static and dynamic performance of SEIG is greatly necessary. This paper contains two sections; the first section deals with the simplified parallel modeling of SEIG by using quadratic equation, which is simpler than non-linear equation and evaluation of minimum capacitance value to fulfill the voltage build-up process. The Second section elaborates the result and discussion. II. EXPANDED PARALLEL MODEL OF SEIG A schematic diagram of stand-alone SEIG for a wind energy conversion system (WECS) is shown in Fig. 1. For transient analysis of induction generator, Park\u2019s transformation model [14] is drawn in Fig. 2 considering saturation effect, which is a non-linear relationship between the magnetizing current and the magnetizing inductance. 978-1-4799-7455-9/15/$31.00 \u00a92015 IEEE The state-space equation of the SEIG is given in Equation (1) .The per phase equivalent circuit, as presented in [3], is a series equivalent circuit. It is simplified to parallel equivalent circuit as given in Fig. 3 to calculate the reactive power required for SEIG" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000914_20130825-4-us-2038.00108-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000914_20130825-4-us-2038.00108-Figure9-1.png", + "caption": "Figure 9. Schematic diagram of a machine link and its dynamic parameters.", + "texts": [ + " The next step is to determine the static and dynamic forces, torques, energy and power consumption from the simulation model. Static torques are calculated with use of the Jacobian, J and its transverse, JT, as follows from Craig (1989): [ ] [ ] (11) The evaluation of the dynamic forces and torques need an advanced computational element, based on manipulator dynamics, summarized as follows. The first step is to apply Newton\u2019s and Euler\u2019s laws to each machine link. For any given link, the notations are shown schematically in Fig. 9. Newton\u2019s law: (12) Euler\u2019s Law: \u0307 (13) Re-map and : (14) \u0307 (15) Re-map , , and \u0307 : (16) (17) \u0307 \u0307 \u0307 (18) Accordingly, the force and torque due to linear and rotational inertia of link i read: (19) \u0307 (20) The mass of each link of the Hitachi 3500 excavator was accessed from the data of the machine. The inertia matrices were calculated from the geometry and the links, simplifying the shape to be prismatic, symmetrical to the center of gravity. While the forces from the weight and inertial components of the machine can be accurately calculated with the presented model from catalog data, the digging force on the bucket is dependent on many factors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002831_ls.1497-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002831_ls.1497-Figure3-1.png", + "caption": "FIGURE 3 Bearing model\u2014mesh details [Color figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " Hence, Equation (11) is used for prediction of the local wear depth; dh ds = kd:p, \u00f012\u00de where k/H is replaced by kD, the dimensional wear rate and the total wear depth, h, for every contact point between the surfaces is formulated from h= \u00f0 kD:p:ds \u00f013\u00de In the current analysis, bearing geometry was made through CAD approach, and the CFD-FSI analysis was carried out with ANSYS\u2013Workbench platform. The bearing is divided into two faces and four edges for meshing, and eight-node hexahedron structure meshes were used as shown in Figure 3. The total number of elements used for the simulation is 46 566, and total nodes are 213 527. In the analysis, as the load assumed to be constant, the eccentricity ratio depends on pressure equilibrium over the journal surface. The mesh is generated for different values of eccentricity ratios, and a mesh quality around 0.50 was maintained for all generated elements. In the current work, the applied boundary conditions were pressure outlet with a gauge pressure of 0 Pa and the calculated lubricant flow rate at inlet zone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001218_tasc.2013.2284021-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001218_tasc.2013.2284021-Figure4-1.png", + "caption": "Fig. 4. The photograph of the PM and HTS bulk annuli. (a) PM, (b) the top view of HTS bulk annuli used in the experiments.", + "texts": [ + " 3 shows the measuring method of levitation forces in the axial and radial directions. The HTS bulks were magnetized by PM with FC method at liquid nitrogen temperature, and PM was levitating between HTS bulks at the same time. Then, we measured the levitation force against the displacement from the levitating position by using a load cell. Here, we defined the restoring force as a levitation force, and we measured not only the levitation force Fz in the axial direction but also Fx in the radial direction. Fig. 4 shows the photograph of the PM and HTS bulk annuli. We used 2 PMs in experiment, the neodymium magnets with ID 10.8 mm, OD 64 mm and thickness 13 mm were used as PM, and their maximum Bz is 0.324 T. In 1 unit system, PM-A was used, and PM-A was used for upper part and PM-B was used for lower part in 2 PM + 3 bulks system as shown in Fig. 2. Fig. 5 shows the measured magnetic flux density Bz along the radial position of 1 mm upper part of 2 PMs, the measured line was shown in Fig. 4(a). In addition, GdBa2Cu3Ox (GdBCO) oxide superconducting bulks with ID 20 mm, OD 60 mm and various thickness (5, 15 and 20 mm) made by quench-meltgrowth (QMG) process were used. With this experimental method, we also measured the levitation force when the shape of PM and material of the rotor shaft were changed. In real applications, the levitation force in the axial and radial directions are very important. Fig. 6 shows the measured Bz along the axial direction of 1 unit system and 2 PM + 3 bulks system before and after field trapping process by FC method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003799_s40313-020-00664-y-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003799_s40313-020-00664-y-Figure2-1.png", + "caption": "Fig. 2 Dump truck lift model", + "texts": [ + " Where m1 is the unsprung mass of the vehicle; m2 is the sprung mass of the frame; m3 is the mass of the cargo box and cargo; A1, A2 and A3 are the roll centers of m1, m2 and m3; \u03d5u , \u03d5s1 and \u03d5s2 are the roll angles of m1, m2 and m3; G1, G2 and G3 are the gravity of each part of the vehicle; \u03b2 is the lateral slope angle of pavement; Y0 is the initial lateral offset of the centroid of m3; d is the wheelbase; hu is the roll length of m1; hr1 is the initial height of the roll center A2; ul and ur are the active suspensions control forces; Fl and Fr are vertical loads acting on the left and right wheels; kt is the vertical stiffness of a single tire; k1 and k2 are the stiffness of the suspension spring and the equivalent stiffness of the box when lifting; c1 and c2 are the equivalent damping coefficients of the suspension and cargo. The target dump truck is set to the front cylinder straight push lift mode, the lower end of the cylinder is connected to the frame through the spherical sub-hinge, and the upper end is connected to the cargo box through the rotary sub-hinge. The vehicle lift model when the road surface is a horizontal plane and the vehicle does not roll is shown in Fig. 2. During the lifting process, the cargo box is subjected to gravity G3, the cylinder lifting force F and the supporting force FN at the hinge point of the container frame. Where L1 is the distance between point C and point O; L2 is the distance between point C and point A; L3 is the distance from point A to point O; b1 and b2 are the vertical distance and horizontal distance between the centroid of the cargo box to the line segment OC and point O; o1 is the angle between the line segment OC and the centroid of the cargo box to the point O when the cargo box is not lifted; o2 is also the angle between the line segment OC and the line segment OB when the cargo box is not lifted; \u03c8 is the angle between the line segment AC and the line segment AO; h is the height that the centroid of the cargo box rise vertically relative to the horizontal road surface; \u03c3 is the lifting angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003308_s40430-020-02488-y-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003308_s40430-020-02488-y-Figure7-1.png", + "caption": "Fig. 7 Logarithmic spiral bevel gear based on Boolean difference", + "texts": [ + " [34] established a precise helix angle conical logarithmic spiral, which was used as the toothed line of the LSBG. The tooth profile of the LSBG was built by adopting the exact involute and arc transition, and the first cogging was afterward created by accurately sweeping along the guide line. The cogging was later arrayed, so that the three-dimensional model of the LSBG was accurately modeled by the Boolean difference between the face cone and the tooth groove. This method solves a problem that the root cone does not intersect with the gear tooth in the sum of the root cone and the tooth, as shown in Fig.\u00a07. The left and middle sides of Fig.\u00a07 show that the cogging entity is evenly distributed on the solid model of the face cone. The right side of Fig.\u00a07 shows the three-dimensional model of the LSBG, which is obtained by the Boolean difference of the model on the left. The mathematical model, based on the cutting principle method, has been developed into basic technology for computer-integrated methods to design, manufacture and analysis LSBGs. Since the mathematical model follows the manufacturing principle of special machine, it closely matches the actual gear. Although complicated, the cutting principle method is a clear spatial transformation process that yields a universal mathematical model adopted by most existing cutting systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003363_j.mechmachtheory.2020.104027-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003363_j.mechmachtheory.2020.104027-Figure4-1.png", + "caption": "Fig. 4. Pulley acceleration measurement devices and noise measurement method [2] .", + "texts": [ + " (1) , it can be seen that chordal displacement z increases when the pulley applied radius R is small or when chain pitch l p is large. In addition, when the input rotation speed is N in (rpm), the noise of a chain-type CVT increases at frequency f 1 st , caused by pitch l p as shown in Eq. (2) [2] . f 1 st = 2 \u03c0R in l p N in 60 (2) To indirectly identify the behavior of a CVT chain that causes noise, the acceleration of the pulley was measured under conditions that cause the sound pressure at frequency f 1 st to increase [2] . Fig. 4 shows the pulley acceleration measurement devices. The input and output pulleys were mounted with two 3-axis accelerometers (356B21 made by PCB) at 180 \u00b0 symmetrical positions on the back. The mounting position radius was 54 mm, which is the pulley\u2019s applied radius when the CVT ratio \u03b3 is 1.0. The obtained data was recorded using a data logger (wireless type) manufactured by i-NEAT for measuring acceleration. The pulley acceleration was measured under the condition that generated the largest chain noise (frequency f 1 st ) at a CVT ratio \u03b3 of 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000954_0954406213516305-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000954_0954406213516305-Figure16-1.png", + "caption": "Figure 16. The accelerometers adsorbed on the bearing housing.", + "texts": [ + " To further validate the usefulness, the proposed method is applied to analyze the experimental vibration signal of a bearing test rig. The facilities of the bearing test rig are shown in Figure 13, and its schematic is displayed in Figure 14. The REB with inner race and rolling-element defects using wire-saw cutting with the crack-fault dimension of 0.2mm width and 0.1mm depth are depicted in Figure 15. Then, it was mounted on the right of the test rig. Vibration data were acquired using accelerometers, which were adsorbed on the surface of bearing housing with magnetic bases as displayed in Figure 16. The vibration signal was acquired using a 4-channel at Gebze Yuksek Teknoloji Enstitu on May 11, 2014pic.sagepub.comDownloaded from ENET-9163 data-acquisition card from NI while the bearing test rig was operating. The REB type is NJ206EM, and the specification is illustrated in Table 1. Vibration signal was collected at a sampling rate of 10 kHz, and the rotational speed is 262 r/min. The characteristic frequency of inner-race defect fi and rolling-element defect fr can be calculated at 33.94Hz and 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003779_icem49940.2020.9270757-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003779_icem49940.2020.9270757-Figure6-1.png", + "caption": "Fig. 6. Temperature distribution with a wave cooling duct on a) duct and windings and laminations at inlet flow rate 10 l/min and temperature 65oC and motor drive point (1100 rpm, 101 Nm), and b) temperature in the stator and housing along a radial line that starts in a tooth and end in the coolant when the power loss is 1.5 kW.", + "texts": [ + " The lamination thermal conductivity is considered anisotropic with a higher value along the laminar sheets and a lower value through the stack. Furthermore, the cooling fluid is here simplified to be pure water although it normally also contains ethylene glycol (in a 50%/50% water-glycol mix). The temperature dependence of the coolants material properties are also ignored, as they are assumed constant. The effect of this simplification is further investigated in [4]. The resulting temperatures achieved when having a wave cooling duct in the casing is shown in Fig. 6. The 2- way coupled model (Fluent) is used in Fig. 6a, whereas the non-coupled model (COMSOL) is used in Fig. 6b as a comparison. For the latter in Fig. 6b, the temperature along a radial line that starts in the stator core between the slots and ends in the coolant is shown. There is a 3 to 6 degrees difference between the aluminum and the coolant, representing a thermal resistance. The temperature distribution in the wave cooling duct for different flow rates (2.5, 5, 10 l/min), when using the 2-way coupled model is shown in Fig. 7 (for the same loss value). It can be seen that the temperature is rather evenly distributed for the high flow rate of 10 l/min whereas there is a more distinct difference between outlet and inlet temperature for the low flow rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003739_0954405420971128-Figure9-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003739_0954405420971128-Figure9-1.png", + "caption": "Figure 9. Prototype and finite element prediction of the (a) \u2018Rooster\u2019 and (b) \u2018Eagle\u2019 coins fabricated from additive manufactured blanks at the end of stroke.", + "texts": [ + " The other colours evolving from light red to dark blue correspond to increasing values of pressure applied by the dies. As seen, the centre of both coins remains free of contact up to the end of stroke due to the recesses of 0.15mm existing in both die surfaces (Figure 5). In contrast, the outer region plastically deforms to fill the intricate details of lettering and other reliefs, and to adjust the diameter of the coin to the geometry of the collar. At the end of stroke, the finite element predicted geometries of the prototype coins are very close to the actual ones (Figure 9). The second topic to be addressed is related to material flow at the centre of the coin blanks containing the complex and intricate contoured holes. Figure 10 shows a top view of the normalised velocity field v= vxy v0j j where vxy is the resultant xy-velocity of the material, and v0 is the velocity of the upper drive tool plate (upper die). As seen, if there is no contact between the centre recess of the die and the coin blank (Figure 10(a)), the deformation of the rooster profile is negligible and no corrective actions need to be taken" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002401_aim.2016.7576929-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002401_aim.2016.7576929-Figure3-1.png", + "caption": "Fig. 3. A schematic representation of the kinematic bicycle model. \u03c6 is the front wheel\u2019s steering angle, l is the length of the bicycle measured from the front wheel\u2019s center to the back wheel\u2019s center, and r is the radius of the circular trajectory of the bicycle\u2019s back wheel. \u03b8 is the angle of rotation between the horizontal coordinate and the bicycle. The bicycle\u2019s back wheel center is attached to the needle tip.", + "texts": [ + " For finding an optimal rotation depth such that the needle tip reaches a desired target location, a model for predicting the path, that the needle tip will follow is needed. In related work, commonly a kinematic model of a bicycle is used, which was first introduced and adapted for the purpose of modeling multi-bend needle tip path trajectories by Park et al. [10] and Webster et al. [11]. Since then, the kinematic bicycle model has become the most commonly used method for needle path planning. A schematic of the bicycle model is shown in Fig. 3. The kinematic equation for a bicycle model in Euclidean space is [12] z\u0307 y\u0307 \u03b8\u0307 \u03c6\u0307 = cos\u03b8 sin\u03b8 tan\u03c6/l 0 v+ 0 0 0 1 \u03c9 (6) where v is the bicycle\u2019s translational velocity and hence the needle insertion velocity, \u03c6 is the front wheel\u2019s steering angle and \u03c9 is the angular velocity of the front wheel\u2019s steering angle and therefore \u03c6\u0307 = \u03c9 . \u03b8 is the angle between the horizontal axis and the bicycle. z is the horizontal and y the vertical position of the bicycle back wheel and thus the position of the needle tip" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003065_s40684-020-00217-3-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003065_s40684-020-00217-3-Figure3-1.png", + "caption": "Fig. 3 Model used in the analysis", + "texts": [ + " Therefore, it is considered that the proper preheating temperature should exceed 400 \u2103 for LAMill. In this study, the proper preheating temperature of the material surface was determined to be 600 \u2103 considering the tensile strength curve. The analysis model was modified by referring to the various blade models actually used in industry due to the limited Table 1 Physical properties of the titanium alloy Density ( kg\u2215m3) Young\u2019s modulus (GPa) Poisson\u2019s ratio 4,430 113 0.34 1 3 distance traveled by the laser module in the currently installed five-axis hybrid machine tool. Figure\u00a03 shows the analysis model used in this study. The mesh method was applied to a hexagonal mesh (Quad/Tri) and 137,338 nodes and 40,314 elements were created for the analysis model. The mesh size of the bottom part of the square onto which the jig was fixed and the back side of blade was 1.0 mm. The mesh sizes of all faces except for the back side of the blade were 0.5 mm. In addition, the mesh size of the laser spot zone was 0.3 mm. The feed rate of the circular laser heat source 3 mm in size was 100 mm/min considering the actual machining speed, and a natural convection condition of 5 W/m2 \u00b0C was applied to the analysis model as a cooling condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003756_1350650120966894-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003756_1350650120966894-Figure8-1.png", + "caption": "Figure 8. Load on discretized point.", + "texts": [ + " The major axis of ellipse is much larger than the minor axis of ellipse and then it is still taken as con- tact line, which is described in Figure 7. In the process of solution, the contact line is discretized into n points (Figure 5). Coordinate of these discretized points can be obtained and then the length of contact line can be calculated. Calculation of load per unit contact line The contact line (major axis of ellipse) is discretized into n points (Figure 5). By the solution of LTCA, the load on discretized point (Figure 8) can be obtained, respectively. The load total of all discretized point constitutes the load on contact line (equation (1)). When loads on all contact lines are calculated, the load distribution on the tooth surface can be obtained (Figure 9). The load per unit contact line can be obtained by dividing the load on contact line by the length of contact line, which is described in equation (2). pi\u00bc Xn i\u00bc1 pin (1) Wi \u00bc pi ai (2) Calculation of comprehensive curvature radius for contact position It has been pointed out earlier that the contact line (major axis of ellipse) is discretized into n points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002149_0954410016654182-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002149_0954410016654182-Figure4-1.png", + "caption": "Figure 4. The chattering phenomenon. (a) Ideal sliding mode motion and (b) actual sliding mode motion.", + "texts": [ + " Here, the sliding mode surface is s \u00bc l\u00fe d dt e \u00bc le\u00fe _e \u00f07\u00de where l4 0 is the sliding mode surface coefficient; e and _e are the position error and velocity error, respectively. If one selects the sign function as the sliding mode control strategy, the state trajectory will be enforced to slide to the equilibrium point along the sliding mode surface. Although the switch control strategy has strong robustness and it is not sensitive to the changes of the system parameters and the external disturbance, an insurmountable chattering phenomenon will appear due to its natural discontinuous control characteristics, as shown in Figure 4. The chattering phenomenon may damage the actuators and deteriorate the control performance. Therefore, signum function sign\u00f0s\u00de should be softened using the hyperbolic tangent function tanh\u00f0 s\u00de to alleviate chattering. The curves of the hyperbolic tangent function with different are shown in Figure 5. tanh\u00f0 s\u00de \u00bc e s e s e s \u00fe e s , 4 0 \u00f08\u00de According to Figure 5, the curve of the hyperbolic tangent function is smooth and continuous at the switching point s \u00bc 0, which reduces the chattering greatly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001289_03043797.2014.1001814-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001289_03043797.2014.1001814-Figure4-1.png", + "caption": "Figure 4. (a) Electric slow cooker in the market, (b) a CAD rendering of assembly from one collaborative team.", + "texts": [ + " Collaborative project #3: manufacturing process development In any design firm, designing a new product requires engineers to carry out \u2018benchmarking\u2019 with best-in-class companies. Benchmarking involves studying similar or competitive products in the markets. This involves dis-assembly of all components, examining how parts were produced. After learning from the benchmarking exercise, the engineers are expected to come up with the best designs/products that have better attributes. Through this project, students had an opportunity to simulate the real tasks of the manufacturing engineers in the real world. An electric slow cooker was chosen for this project as shown in Figure 4(a). 4.1.3.1. Project requirements. (a) All components of the product must be disassembled, measured and converted into CAD models, (b) the components can be fabricated by any combination of the manufacturing processes, and (c) the deliverable of this project was a final team report. The report had to include, but not limit to, (1) abstract, (2) introduction, (3) description of all components and CAD models, (4) the proposed manufacturing processes and sequences for fabricating all components, (5) explaining what machines, equipment, and operator skills that were needed for the proposed processes, and justifying the proposed manufacturing process/sequence in terms of time and cost, and (6) Appendix that includes all supportive documents such as emails, meeting minutes, and responsibilities of each member and so on. Figure 4(b) shows a CAD rendering of the product prepared by one of the collaborative teams. As discussed in the introduction section, the major goal of this study was to develop a manufacturing course that will address some of the challenges in manufacturing education. Three specific goals were set to be achieved through this course: (a) to increase student\u2019s knowledge of manufacturing and high technology, (b) to increase students\u2019 preparedness and confidence in effective communication, and (c) to increase students\u2019 interest in pursuing additional academic studies and/or a career path in manufacturing and high technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002451_cacs.2015.7378369-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002451_cacs.2015.7378369-Figure4-1.png", + "caption": "Figure 4. Illustration of the current reconstruction problem at the first sampled time: (a) Reference voltage vector in Sector I boundary region. (b) Corresponding three-phase PWM pattern. ( c) Voltage vectors for switching (proposed). (d) Corresponding three-phase PWM pattern (proposed).", + "texts": [ + " This study mainly focuses on adjusting the switching scheme without adding extra circuits, and keeps the PWM duration a double-sided balance. Figs. 4-6 illustrate the sampling problems and the switching adjustment technique. The double-sided modulation is divided into a sample phase and an acquisition phase. This technique ensures a low-noise sampling, and an average value of the phase current is measured at the same time. In order to facilitate the explanation of the proposed technique, consider the reference voltage III Sector I as an example for the following cases. Fig. 4(a) shows the area where the problem appears and Fig. 4(b) presents an example of the corresponding PWM signals. In this case, two phase currents need to be collected at the instants of V2 and V; . One of the output currents is always correctly sampled because of the sufficient amount of time. The problem occurs at the first sampled time tsampl . The distinguishing feature of the PWM signal waveform is two-long and one-short. When Vre/ is close to V; , it implies that half of the V2 duration is possibly less than Tmin and therefore cannot ensure the current measurement validity at instant tsamPI ", + "T; = 2T;run - T; is inserted in the zero voltage vector to keep the average voltage vector unchanged. The purpose is to fulfill the requirement of the minimum time Tmm for the half PWM interval. Although the amount of switching is increased by one time, the pulse width of each inverter switch is still the same. The adjusting scheme does not change the duty ratios of the switching waveforms, and therefore, the average reference voltage remains the same during a switching interval. The operating principles of the proposed method are shown in Fig. 4(c) and (d). The method ensures the PWM output signals are divided into two parts equally and keep the waveform symmetrical. Furthermore, the add-on voltage vector \ufffd/ ensures the current measurement is valid during the two switching states and the sampling can be completed at the left side within one PWM interval. As illustrated in the previous section, Fig. Sea) and (b) presents a similar problem and its corresponding PWM signals, respectively. The problem occurs at the second sampled time tSGmp2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002581_icumt.2016.7765227-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002581_icumt.2016.7765227-Figure1-1.png", + "caption": "Fig. 1. Twin Rotor MIMO System", + "texts": [ + " In this paper, the same control technique is applied to a linearized model of TRMS and implemented to the laboratory platform. The proposed approach is considerate of both desired quality criteria and required robustness to undesired influences of parametric uncertainties and external disturbances. Presented experimental results are confirmation of that. II. MODEL DESCRIPTION Twin Rotor MIMO System laboratory platform is a helicopter-like nonlinear system with strong cross reactions and independent control for each degree of freedom. General view of TRMS is presented in Fig.1. Plant model analysis in sections II, III was financially supported by Government of Russian Federation, Grant 074-U01. Synthesis of control system in section V was supported by the Ministry of Education and Science of Russian Federation (Project 14.Z50.31.0031). Formulation of the Theorem in section V was supported by the Russian Federation President Grant (No. 14. W01.16.6325-MD (No. MD-6325.2016.8)). Control system approbation on laboratory platform was supported by the Russian Federation President Grant \u211614" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003194_s12206-020-0503-y-Figure14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003194_s12206-020-0503-y-Figure14-1.png", + "caption": "Fig. 14. Established testing platform.", + "texts": [ + " The rotating frequency after rotor drops onto the traditional TDB decreases significantly fast, which is mainly due to the continuous collisions and friction between the rotor and the inner race. Fig. 11. Collision force after rotor drop onto traditional TDB. 0 Time t/s 0.7 1.4 2.1 2.8 3.5 0.04 0.08 0.12 0.16 0.20 0.16 0.17 0.18 0.19 0.20 0 50 100 150 200 t/s Traditional catcher bearing New type catcher bearing Fig. 12. Average collision force after rotor drops onto different types of TDB. The rotor drop test platform is established, as shown in Fig. 14, to verify the working performance of the new-type active TDB. On the basis of the original MBC control system, the new-type active TDB controller is added and developed based on DSP28335. The DSP\u2019s A/D module is used to sample the output signal of displacement sensors. When the rotor is in the set phase, the active TDB controller will pull down the driving signal of the power switch tube in the MBC power amplifier corresponding to the left end. The current in the AMB coil enters into the natural subsequent flow stage, and the support forces of AMB applied on the left rotor are sheared" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001001_j.protcy.2014.08.013-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001001_j.protcy.2014.08.013-Figure2-1.png", + "caption": "Fig. 2: Leg model", + "texts": [ + " The actual reachable workspace of each leg is a sector of an annulus which also overlaps with the adjacent area as per the kinematic property of a multi legged robot. A rectangular region is chosen for every leg from the annulus area to avoid the interference. Reachability is compromised in such case. The stroke pitch is assumed to be the same as the length of a working area, Rx .That is, and as shown in the fig. 1 working areas of the legs are adjacent to one another. The leg model used in the hexapod as shown in the fig. 2 is composed of three rigid links. The body and the three links are connected to one another by help of active revolute joints. The joint at the main actuator is joint one, the joint at lifting actuator is joint two, and the joint at knee actuator is joint three. First joint axis is perpendicular to the plane of the body and the other two are parallel to the longitudinal direction. Hence the robot has three degrees of freedom and walking can take place in any direction. \u03b81, \u03b82 and \u03b83 are the respective joint angles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000656_s106836661405016x-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000656_s106836661405016x-Figure4-1.png", + "caption": "Fig. 4. Calculated trajectory of (a) center of the shaft and (b) calculated pressures in bearing of GTK 10I gas turbine compressor.", + "texts": [ + " Along with other testing calculations, these results allow one to draw a conclusion about the JOURNAL OF FRICTION AND WEAR Vol. 35 No. 5 2014 HYDRODYNAMIC ANALYSIS OF FRICTION BEARINGS 399 applicability of the described algorithms for hydrody namic calculations of real slide bearings. SOFTWARE IMPLEMENTATION Developed methodology for calculating hydrody namic characteristics of bearings of gas turbine com pressor GTK10 I, crosshead bearings of marine die sel, high pressure pump station UNP55 250, opposite compressor and other aggregates. Figure 4 shows the shaft trajectory and pressure distribution in the bearing of a GTK10 I gas turbine compressor with elliptic housing. Figure 4b illustrates two oil wedges on the top and bottom insets, which promote the decrement of shaft vibration. Presented plots and diagrams were plotted in wide spread software packages such as MATHCAD. The further development of works was directed to the extension of functionality (consideration of more impacting factors) and upgrading of the interface. Presently, several versions of Bearing Builder Finite Element Method (BBFEM) programs have been developed that differ in their functional possibilities by the Department of the Dynamics and Strength of Machines at Bryansk State Technical University", + " After discretization, the thickness of the layer within each triangular FE is defined by the bilinear interpolation of the real thicknesses (Fig. 8). In these cases, it makes sense to apply mesh with refined FE in the places of deviations. Figure 9a shows an example of the generation of mesh in the presence of lubricant grooves on the bearing surface using the Ruppert algo 400 JOURNAL OF FRICTION AND WEAR Vol. 35 No. 5 2014 ZERNIN et al. rithm [9]. The surface relief is visualized in order to control the preparation of data (Fig. 9b). As in the case of the ellipticity of the bearing hous ing (Fig. 4), the methodology allows one to model other types of constructive deviations from cylindric ity, e.g., the system of lubrication grooves of the cross head diesel bearing (Fig. 10) and the calculation of pressure (Fig. 11). In order to model irregularities of the path by which the shaft surface passes near each irregularity is discretized by a fine mesh of FEs (Fig. 12a). At each moment of time, the hydrodynamic problem of taking into account the results of the previous step based on the angle of shaft rotation, e" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001169_gt2014-26891-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001169_gt2014-26891-Figure2-1.png", + "caption": "Figure 2: A Timoshenko Beam segment", + "texts": [ + " The Newmark-\u03b2 implicit time integration scheme along with Newton-Raphson method is used to derive steady state solution of non-linear system of equations of motion. Considering different types of misalignments in succession and combination, the dynamic response is characterized by the presence of 1N, 2N and, other super or sub harmonics. The finite element model of each flexible rotor is developed using circular Timoshenko beam finite elements to include shear deformation. The gyroscopic moments, rotary inertia, and damping for bearing and shaft material are also accounted for. The XZ and YZ analogous motion planes [13], as shown in Fig. 2 are assumed to be uncoupled. The equations of motion of a rotating Timoshenko beam element, in two perpendicular motion planes are 2 2 2 2 ( ) ( , ) ( ) ( , ) S x z S y z x xAG p Z t m Z Z t y yAG p Z t m Z Z t \u03c1 \u03b1 \u03c1 \u03b2 \u2202 \u2202 \u2202\u23a1 \u23a4\u2212 + =\u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6 \u2202 \u2202 \u2202\u23a1 \u23a4\u2212 + =\u23a2 \u23a5\u2202 \u2202 \u2202\u23a3 \u23a6 2 2 2 2 ( ) ( ) ( ) ( ) S T r z S T r z xEI AG J J Z Z Z t t yEI AG J J Z Z Z t t \u03b1 \u03b1 \u03b2\u03c1 \u03b1 \u03c9 \u03b2 \u03b2 \u03b1\u03c1 \u03b2 \u03c9 \u2202 \u2202 \u2202 \u2202 \u2202 + \u2212 = + \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 + \u2212 = \u2212 \u2202 \u2202 \u2202 \u2202 \u2202 (1) 3 Copyright \u00a9 2014 by ASME Downloaded From: http://proceedings.asmedigitalcollection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000630_ilt-11-2011-0103-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000630_ilt-11-2011-0103-Figure2-1.png", + "caption": "Figure 2 Bearing assembly cross-section", + "texts": [ + " Design and fast verification of pocket elliptical journal bearings Fabrizio Stefani Industrial Lubrication and Tribology Volume 66 \u00b7 Number 3 \u00b7 2014 \u00b7 393\u2013401 In Figure 1, if O1 and O2 denote, respectively, the centers of the bottom and top pads of the \u201celliptical\u201d part of the bearing and Oc is the center of the cylindrical pocket, the center of the turning operation (the location of spindle center) is coincident with both O1 and O2. The point O, which is set in the middle of the distance POc, is the origin of the reference system used to determine the offset s OO1 OO2 of the lathe spindle axis. The shim thickness is divided in two parts OcO1 and PO1 measured by the parameters e1 and e2, which are referred to as the eccentricities of the bottom and top pads, respectively. Such designation is explained by Figure 2, where the final bearing profile is shown. After the second machining, the two halves are joined together so that the top pad drops until its center O2 moves below Oc at a distance e2, while O1 remains in the same location, at distance e1 from Oc. The radius of the surface machined during the second operation is: Rp Rc t R cp (2) where t is the cut depth and cp the clearance of the elliptical bearing pads. The eccentricity e1 is an important parameter, as it rules the bottom pad profile that determines the load-carrying capacity of the bearing in nominal conditions (the bottom pad preload m1 e1/cp)", + " Figure 1 allows one to calculate e1 by means of simple geometrical considerations, as follows: e1 O1Oc O1H OcH Rp 2 Rc 2 sin2 Rc cos (3) or to express e1 as a function of manufacturing parameters by means of the following equation: e1 s d 2 (4) Similarly, the top pad eccentricity is given by: e2 d 2 s (5) As the same bottom pad eccentricity e1 can be obtained by means of several (infinite) couples of s and d according to equation (4), different bearing designs exist characterized by the same bottom lobe profile. Two limit cases, which will be denoted with the subscripts A and B, can be obtained by assuming e2 0 and e2 t, respectively. For the purpose of the implementation of the numerical code, the bearing clearance distribution must be calculated. Figure 2, by showing the journal profile when its center overlaps Oc, illustrates the bearing clearance parameters. At the end of machining, the elliptical pad clearance cp, in agreement with equation (2), is equal to the horizontal clearance cH of the bearing, if higher-order infinitesimals (the lengths AB and A\u2019B\u2019) are neglected. Let cv1 and cv2 be the vertical clearances of bottom and top pad, respectively. By comparing Figure 1 to Figure 2 and by using equations (1) and (2), the following assembly constraint equations can be formulated: cv1 cc (6) cV2 cp e2 cc t e2 (7) For design A (e2 0), the resulting vertical clearance on the bottom and top pads are cV1A cc and cv2A cp cc t, respectively, i.e. clearance is higher in the top pad than in the bottom pad. In nominal working conditions, such additional clearance does not cause higher side loss, as the oil on the top pad, due to cavitation, flows only in the circumferential direction (Couette flow)", + " On the contrary, in transient operations, such as during the start-up aided by hydrostatic Design and fast verification of pocket elliptical journal bearings Fabrizio Stefani Industrial Lubrication and Tribology Volume 66 \u00b7 Number 3 \u00b7 2014 \u00b7 393\u2013401 lifting, the formation of an active (pressurized) film in the top pad may give rise to high lubricant side loss. Consequently, some manufacturers resort to design B (e2 t). In such a case, the vertical clearance of both the bush halves is the same [cv1B cv2B cc, in agreement with equations (6) and (7)], as the edge of the top pad is tangent to the cylindrical profile machined by the first turning operation (Figure 2 with tv2 0). Independently of the design solution, with reference to Figure 2, the cut depth at the bottom and top pad center (on the load line), tv1 and tv2, respectively, in agreement with equations (2) and (7) are defined by: tv1 e1 t (8) tv2 cv2 cc t e2 (9) The second turning operation may be performed either on the whole length L of the bearing land (in the axial direction) or on a part of it, symmetrically with reference to the load line. In the latter case, schematized in Figure 3, on each atmospheric side of the bearing, a cylindrical surface with axial length Lb and uniform radial clearance cb cc is left" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003753_jestpe.2020.3037942-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003753_jestpe.2020.3037942-Figure12-1.png", + "caption": "Fig. 12. The 6s13r CP-SFPM machine Prototype. (a) Stator. (b) Rotor.", + "texts": [ + "225 |7ZPM-Zr|, 29th 0.010 0.008 0.130 |3ZPM+Zr|, 31th 0.110 0.089 1.328 Sum (100%)/ FE predicted (Nm) 6.685/6.754 Flux barrier effect Flux leakage|3ZPM-Zr| 5th Flux leakageFlux barrier effect |5ZPM-Zr| 2nd (a) (b) In order to verify above theoretical analyses, an optimal CP-SFPM prototype with iron bridge design is manufactured. The stator and rotor assemblies are shown in Figs. 12(a) and (b), respectively. The iron bridge design in stator yoke is adopted to enhance mechanical strength, as shown in Fig. 12(a). The measured and 2D FE predicted back-EMFs are shown in Figs. 13(a) and (b), respectively. The comparison of the measured and 2D FE predicted fundamental components of the back-EMFs under different rotor speeds is shown in Fig. 13(c). The measured back-EMF results are slightly lower than the 2D FE predictions, which is mainly due to the end effect and mechanical tolerance are neglected in 2D FE simulation. Besides, the comparison of the measured and 2D FE predicted steady-state torque curves with different phase currents is given in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001150_epe.2013.6632004-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001150_epe.2013.6632004-Figure8-1.png", + "caption": "Fig. 8 : Equivalent scheme of the stator (in the dq frames)", + "texts": [ + " The Joules effect transforms the electric energy in thermal power. The iron losses generate thermal energy by magnetic effect. The friction on the shaft transforms the mechanical energy in thermal energy. a) Electrical -Thermal Description The Joules and friction effect are previously considered in the different models (2) and (5), but these losses are not highlighted with the classical description. In order to highlight the thermal dissipation of Joules and iron effect, the equivalent stator scheme (Fig. 8) can be separated into three parts. Firstly the resistance effect Rs (8), secondly the iron losses (9) and thirdly the inductance Ldq effect (10). An equivalent resistance RI is introduced to represent the iron losses. This resistance value RI depends on the design and the rotation speed \u03a9 [9]. dqsrdq IRV =_ (8) )( _ _ \u2126R V I I idq idq = (9) dqLdqidq eI dt dLV += __ (10) Ldqidqdq III __ += (11) idqrdqdq VVV __ += (12) The EMR of the stator is composed of 5 elements (the resistance, the inductance, the iron losses and 2 coupling) (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000846_j.ymssp.2014.11.015-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000846_j.ymssp.2014.11.015-Figure1-1.png", + "caption": "Fig. 1. Example case: a mechanical system with an elliptic cam and follower supported by a lumped spring (ks).", + "texts": [ + " In contrast, a cam\u2013follower mechanism with rotational sliding contact (with no impacts) is used to experimentally determine \u03bc in this study. Since \u03bc cannot be directly measured from vibration experiments, an analogous contact mechanics model [23] is developed to aid the process. The goal is to vary the surface roughness, lubrication film thickness, contact pressure and velocities at contact (sliding and entrainment). The proposed system could then be utilized to simulate the contact conditions seen in drum brakes and geared systems. Fig. 1 shows the mechanical system with an elliptic cam (with semi-major and minor axes as a and b, respectively). Though the kinematics of combined rolling\u2013sliding contact systems are complex compared to systems with pure sliding contact, this is one of the simplest systems which one can devise to measure the coefficient of friction in such systems which would allow controlled measurements of the reaction forces and system acceleration. The cam is pivoted at E along its major axis with a radial run out, e, from its centroid (Gc, with subscript c denoting cam)", + " The equation of the elliptic cam is given by the following, where r is the radial distance from Gc to any point on the circumference of the cam, and \u0394 is the polar angle of that point, r \u0394 \u00bc abffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a sin \u0394 2\u00fe b cos \u0394 2q : \u00f01\u00de The cam is in a point contact (at Oc) with the follower (at Ob, with subscript b denoting follower), which consists of a thin cylindrical dowel pin (of radius rd) attached to a bar (of length lb) of square cross-section (of width wb). The center of gravity of the follower lies at Gb at a distance of lg from the pivot point P (using roller bearings) which is at dy distance above the ground. The follower is supported by a linear spring (ks) along the vertical direction (e\u0302y), which is at a distance of dx from P as shown in Fig. 1. The angular motion of the follower is given by \u03b1(t) in the clockwise direction from the e\u0302x axis; it is also the only dynamic degree-of-freedom of the system. The contact mechanics at O between the cam and the follower is represented by non-linear contact stiffness (k\u03bb) and viscous damping (c\u03bb) elements. Viscous damping is valid in this study since the system is designed to not lose the contact at any point of time, and the indentation velocity is low (no impacts), and hence the contact damping force is insignificant (compared to contact stiffness force) regardless the damping mechanisms involved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002121_tcst.2016.2572165-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002121_tcst.2016.2572165-Figure1-1.png", + "caption": "Fig. 1. Kinematic representation of a rigid body with respect to an earth-fixed reference frame.", + "texts": [ + " The model consists of six DOF, three rotational DOF to determine its orientation, and three translational DOF to determine its position. In order to quantify the rigid-body motion, we assign an earth-fixed reference frame, XeYe Ze, at a reference location and a body-fixed frame, XY Z , at the CM of the rigid body. Three position variables x , y, and z are required to position the origin of XY Z relative to XeYe Ze, representing the translational motion of the rigid body with re = [x, y, z]T , as shown in Fig. 1. Normally, three consecutive rotations are required to represent the orientation though other methods such as those using quaternions can also be employed. In this paper, we use roll (\u03c6), pitch (\u03b8 ), and yaw (\u03c8) Euler angles to represent XY Z rotations with respect to XeYe Ze through intermediate reference frames X \u2032Y \u2032Z \u2032 and X \u2032\u2032Y \u2032\u2032Z \u2032\u2032, respectively, as shown in Fig. 2. The directions of roll rate (\u03c6\u0307), pitch rate (\u03b8\u0307 ), and yaw rate (\u03c8\u0307) are, respectively, about X , Y \u2032, and Ze. Thus, the pitch rate axis is orthogonal to both the yaw and roll rate axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001771_robio.2015.7419104-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001771_robio.2015.7419104-Figure4-1.png", + "caption": "Figure 4. Patterns of magnetic circuit.", + "texts": [ + " The cam mechanism of the driver enables extension and contraction of the unit and magnet rotation. The structure of the magnetic adhesion mechanism installed in each unit is shown in Fig. 3. There is a permanent magnet in the center of the mechanism, as shown in Fig. 3 (b). This magnet can rotate inside the mechanism and become magnetized in its radial direction. In this way, the direction of its magnetic pole can be switched between the adhesion and roller side yoke. Magnet rotation allows the magnetic adhesion mechanism to switch between two magnetic circuit patterns, as shown in Fig. 4. In Pattern 1, the magnetic pole of the permanent magnet faces the adhesion yokes, which then establish a magnetic circuit through the adsorption surface. In this scenario, the magnet adhesion mechanism produces a large adsorption force. Further, the adhesion yokes are attached to friction plates (see Fig. 3 (a)). When the magnet adhesion mechanism adsorbs onto a surface, this structure imparts a high frictional force to the surface. In Pattern 2, the magnetic pole of the magnet faces the roller side yokes that form a magnetic circuit through the magnetic rollers and adsorption surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure10.17-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure10.17-1.png", + "caption": "Figure 10.17 illustrates the top view of the manipulator in a pose such that the center of the moving platform is vertically below the center of the fixed platform. As seen in the figure, the manipulator has a symmetric configuration such that the fixed platform (0) can be modeled geometrically as an equilateral triangle. The centers of the revolute joints (B1, B2, B3) are located at the corners of this triangle. Considering one of the three legs, say Lk , its upper link is k and its lower parallel links are \u2032 k+3 and \u2032\u2032 k+3. On these parallel links, the centers of the upper spherical joints are C\u2032 k and C\u2032\u2032 k with the mid point Ck between them. Similarly, the centers of the lower spherical joints are A\u2032 k and A\u2032\u2032 k with the mid point Ak between them. The lower spherical joints are shared by the moving platform 7.", + "texts": [], + "surrounding_texts": [ + "Referring to Figures 10.17 and 10.18, the following equations can be written in order to express the orientation of the moving platform (7) together with the end-effector and the location of the tip point (P) with respect to the base frame 0(O) by going through each leg. Position and Velocity Analyses of Parallel Manipulators 405 (a) End-Effector Orientation Equations Through the Legs For k = 1, 2, 3, 7 and the \u201clefty\u201d (left-hand side) legs, C\u0302 = C\u0302(0,7) = C\u0302(0,0k)C\u0302(0k,k)C\u0302(k,k+3)\u2032C\u0302(k+3,7k)\u2032C\u0302(7k,7)\u2032 \u21d2 C\u0302 = [eu\u03033\ud835\udefdk ][eu\u03032\ud835\udf03k ][eu\u03032\ud835\udf03 \u2032 k eu\u03033\ud835\udf19 \u2032 k eu\u03031\ud835\udf13 \u2032 k ][eu\u03031\ud835\udf13 \u2032 k+3 eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3][eu\u03033\ud835\udefdk ]t \u21d2 C\u0302 = eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 e\u2212u\u03033\ud835\udefdk (10.441) For k = 1, 2, 3, 7 and the \u201crighty\u201d (right-hand side) legs, C\u0302 = C\u0302(0,7) = C\u0302(0,0k)C\u0302(0k,k)C\u0302(k,k+3)\u2032\u2032C\u0302(k+3,7k)\u2032\u2032C\u0302(7k,7)\u2032\u2032 \u21d2 C\u0302 = [eu\u03033\ud835\udefdk ][eu\u03032\ud835\udf03k ][eu\u03032\ud835\udf03 \u2032\u2032 k eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031\ud835\udf13 \u2032\u2032 k ][eu\u03031\ud835\udf13 \u2032\u2032 k+3 eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 ][eu\u03033\ud835\udefdk ]t \u21d2 C\u0302 = eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 e\u2212u\u03033\ud835\udefdk (10.442) Note that the Euler angle sequences associated with the upper and lower spherical joints are selected differently (i.e. respectively, as 2-3-1 and 1-3-2) for the sake of kinematic convenience as it will be evidenced in the sequel. Here, it can be said that the major convenience of these sequences is that the indefinite angle pairs {\ud835\udf13 \u2032 k, \ud835\udf13 \u2032 k+3} and {\ud835\udf13 \u2032\u2032 k , \ud835\udf13 \u2032\u2032 k+3} that describe the insignificant spinning motions of the lower legs get combined into single equivalent spin angles. (b) Tip Point Location Equations Through the Legs \u2217 Tip point location equations through the lefty legs for k = 1, 2, 3: p\u20d7 = \u2212\u2192OP = \u2212\u2192OB0 + \u2212\u2212\u2192B0Bk + \u2212\u2212\u2192BkCk + \u2212\u2212\u2212\u2192 CkC\u2032 k + \u2212\u2212\u2212\u2192 C\u2032 kA\u2032 k + \u2212\u2212\u2192 A\u2032 kAk + \u2212\u2212\u2192AkA0 + \u2212\u2212\u2192A0P \u21d2 p\u20d7 = h0u\u20d7(0) 3 + b0u\u20d7(0k) 1 + b1u\u20d7(k) 1 + d0u\u20d7(k) 2 + b2u\u20d7(k+3)\u2032 1 \u2212 d0u\u20d7(7k)\u2032 2 \u2212 b7u\u20d7(7k)\u2032 1 \u2212 h7u\u20d7(7) 3 (10.443) Equation (10.443) can be written as the following matrix equation in the base frame. p = p(0) = h0u(0\u22150) 3 + b0u(0k\u22150) 1 + b1u(k\u22150) 1 + d0u(k\u22150) 2 + b2u(k+3\u22150)\u2032 1 \u2212 d0u(7k\u22150)\u2032 2 \u2212 b7u(7k\u22150)\u2032 1 \u2212 h7u(7\u22150) 3 \u21d2 p = h0u3 + b0C\u0302(0,0k)u1 + b1C\u0302(0,k)u1 + d0C\u0302(0,k)u2 + b2C\u0302(0,k+3)\u2032u1 \u2212 d0C\u0302(0,7k)\u2032u2 \u2212 b7C\u0302(0,7k)\u2032u1 \u2212 h7C\u0302(0,7)u3 \u21d2 p = h0u3 + b0eu\u03033\ud835\udefdk u1 + b1eu\u03033\ud835\udefdk eu\u03032\ud835\udf03k u1 + d0eu\u03033\ud835\udefdk eu\u03032\ud835\udf03k u2 + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031\ud835\udf13 \u2032 k u1 \u2212 d0eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 u2 \u2212 b7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 u1 \u2212 h7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 e\u2212u\u03033\ud835\udefdk u3 \u21d2 p = h0u3 + b0eu\u03033\ud835\udefdk u1 + b1eu\u03033\ud835\udefdk eu\u03032\ud835\udf03k u1 + d0eu\u03033\ud835\udefdk u2 + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k u1 \u2212 d0eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 u2 \u2212 b7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 u1 \u2212 h7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 u3 \u21d2 406 Kinematics of General Spatial Mechanical Systems p = h0u3 + eu\u03033\ud835\udefdk [u1(b0 + b1c\ud835\udf03k) + u2d0 \u2212 u3b1s\ud835\udf03k] + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k u1 \u2212 eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k eu\u03031(\ud835\udf13 \u2032 k+\ud835\udf13 \u2032 k+3)eu\u03033\ud835\udf19 \u2032 k+3 eu\u03032\ud835\udf03 \u2032 k+3 (u1b7 + u2d0 + u3h7) (10.444) Equation (10.444) can also be written more compactly as follows upon inserting the matrix C\u0302 given by Eq. (10.441). p = h0u3 + eu\u03033\ud835\udefdk [u1(b0 + b1c\ud835\udf03k) + u2d0 \u2212 u3b1s\ud835\udf03k] + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032k )eu\u03033\ud835\udf19 \u2032 k u1 \u2212 C\u0302eu\u03033\ud835\udefdk (u1b7 + u2d0 + u3h7) (10.445) \u2217 Tip point location equations through the righty legs for k = 1, 2, 3: p\u20d7 = \u2212\u2192OP = \u2212\u2192OB0 + \u2212\u2212\u2192B0Bk + \u2212\u2212\u2192BkCk + \u2212\u2212\u2212\u2192 CkC\u2032\u2032 k + \u2212\u2212\u2212\u2212\u2192 C\u2032\u2032 k A\u2032\u2032 k + \u2212\u2212\u2212\u2192 A\u2032\u2032 k Ak + \u2212\u2212\u2192AkA0 + \u2212\u2212\u2192A0P \u21d2 p\u20d7 = h0u\u20d7(0) 3 + b0u\u20d7(0k) 1 + b1u\u20d7(k) 1 \u2212 d0u\u20d7(k) 2 + b2u\u20d7(k+3)\u2032\u2032 1 + d0u\u20d7(7k)\u2032\u2032 2 \u2212 b7u\u20d7(7k)\u2032\u2032 1 \u2212 h7u\u20d7(7) 3 (10.446) Equation (10.446) can be written as the following matrix equation in the base frame. p = p(0) = h0u(0\u22150) 3 + b0u(0k\u22150) 1 + b1u(k\u22150) 1 \u2212 d0u(k\u22150) 2 + b2u(k+3\u22150)\u2032\u2032 1 + d0u(7k\u22150)\u2032\u2032 2 \u2212 b7u(7k\u22150)\u2032\u2032 1 \u2212 h7u(7\u22150) 3 \u21d2 p = h0u3 + b0C\u0302(0,0k)u1 + b1C\u0302(0,k)u1 \u2212 d0C\u0302(0,k)u2 + b2C\u0302(0,k+3)\u2032\u2032u1 + d0C\u0302(0,7k)\u2032\u2032u2 \u2212 b7C\u0302(0,7k)\u2032\u2032u1 \u2212 h7C\u0302(0,7)u3 \u21d2 p = h0u3 + b0eu\u03033\ud835\udefdk u1 + b1eu\u03033\ud835\udefdk eu\u03032\ud835\udf03k u1 \u2212 d0eu\u03033\ud835\udefdk eu\u03032\ud835\udf03k u2 + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031\ud835\udf13 \u2032\u2032 k u1 + d0eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 u2 \u2212 b7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 u1 \u2212 h7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 e\u2212u\u03033\ud835\udefdk u3 \u21d2 p = h0u3 + b0eu\u03033\ud835\udefdk u1 + b1eu\u03033\ud835\udefdk eu\u03032\ud835\udf03k u1 \u2212 d0eu\u03033\ud835\udefdk u2 + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k u1 + d0eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 u2 \u2212 b7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 u1 \u2212 h7eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3 u3 \u21d2 p = h0u3 + eu\u03033\ud835\udefdk [u1(b0 + b1c\ud835\udf03k) \u2212 u2d0 \u2212 u3b1s\ud835\udf03k] + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k u1 \u2212 eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k eu\u03031(\ud835\udf13 \u2032\u2032 k +\ud835\udf13 \u2032\u2032 k+3)eu\u03033\ud835\udf19 \u2032\u2032 k+3 eu\u03032\ud835\udf03 \u2032\u2032 k+3(u1b7 \u2212 u2d0 + u3h7) (10.447) Equation (10.447) can also be written more compactly as follows upon inserting the matrix C\u0302 given by Eq. (10.442). p = h0u3 + eu\u03033\ud835\udefdk [u1(b0 + b1c\ud835\udf03k) \u2212 u2d0 \u2212 u3b1s\ud835\udf03k] + b2eu\u03033\ud835\udefdk eu\u03032(\ud835\udf03k+\ud835\udf03\u2032\u2032k )eu\u03033\ud835\udf19 \u2032\u2032 k u1 \u2212 C\u0302eu\u03033\ud835\udefdk (u1b7 \u2212 u2d0 + u3h7) (10.448) Position and Velocity Analyses of Parallel Manipulators 407" + ] + }, + { + "image_filename": "designv11_34_0000338_1077546315586495-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000338_1077546315586495-Figure2-1.png", + "caption": "Figure 2. Modes of the rear twist beam axle.", + "texts": [ + " Polygonal wear mostly occurs under long-term high-speed traveling, and the operating conditions are relatively steady, so the dynamic load of a car body can be ignored in this model, and only translational motion of a car body in the x direction is taken into account. 3. According to static analysis (Wu et al., 2009) and modal analysis of the rear twist beam axle, it is found that there mainly exist three kinds of deformation on the axle when it works: torsional deformation around the n02 coordinate axis and bending deformations around the n01 and n03 coordinate axes, shown in Figure 2. Stiffness of the wheel in the toe-in angle direction and camber angle direction is much greater than that of the rear twist beam axle, so it is assumed that the half rear beam axle is fixed at UNIV CALIFORNIA SANTA BARBARA on July 13, 2015jvc.sagepub.comDownloaded from with the tired wheel at point G, called the rear-axlebody, which connects with the car body by a spherical hinge at point O, using rotational angles , , to reflect deformation in the n01, n02, n03 directions, and parameters of wheel alignment such as the toe-in angle and camber angle will be treated as initial conditions of and " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002050_0954406216648354-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002050_0954406216648354-Figure4-1.png", + "caption": "Figure 4. The processing of curve-face gear: (1) the pitch curve of shaper cutter, (2) the trajectory of shaper cutter and", + "texts": [ + " Because the surface of non-circular is the conjugate tooth surface with shaper cutter, according to the space coordinate transformation theory, the mathematical equation of the tooth surface of non-circular gear 2 in coordinate O2 X2Y2Z2, can be established as follows r \u00f02\u00de 2 \u00f0 s, 1\u00de\u00bc rss\u00bdcos\u00f0 s os\u00de s sin\u00f0 s os\u00de \u00fe L cos\u00f0 1 l\u00de 8>>< >>: 9>>= >>; rss\u00bdsin\u00f0 s os\u00de \u00fe s cos\u00f0 s os\u00de \u00fe L sin\u00f0 1 l\u00de 8>>< >>: 9>>= >>; us 1 2 6666666666666664 3 7777777777777775 \u00f07\u00de The tooth surface of curve-face gear is founded by the involute cylindrical shaper cutter with the external generating method, the shaper cutter is rotating counter-clockwise and the curve-face gear is rotating clockwise in the process of meshing. Figure 4 shows the relationships of the positions and angles of the shaper cutter and curve-face gear. It is shown in Figure 4 that the coordinate system rigidly linked with the frame of the shaper cutter is O1 X1Y1Z1 and the coordinate system rigidly linked with the shaper cutter is O01 X01Y 0 1Z 0 1. The coordinate system rigidly linked with the frame of the trajectory is O2 X2Y2Z2 and the coordinate system rigidly linked with the trajectory is O02 X02Y 0 2Z 0 2. The coordinate system rigidly linked with the frame of the curve-face gear is O3 X3Y3Z3 and the coordinate system rigidly linked with the curve-face gear is O03 X03Y 0 3Z 0 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000762_gt2013-95442-Figure16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000762_gt2013-95442-Figure16-1.png", + "caption": "FIGURE 16. Full section view of the new rotor", + "texts": [ + " There is no real difference between the adapter on the motor\u2019s end and the displacement\u2019s end, except the staggered surface on the motor side as a flow protection geometry for the brush seals, compared to the even running surface on the displacement side. Otherwise both have the same dimensions, with an -compared to the old rotor system- unmodified maximum rotor diameter of 300 mm. To avoid a lifting of the adapters from the rotor due to the high centrifugal load, resulting from the high rotational velocities, a circumferential centring device for each adapter was realised as show in FIGURE 16. FIGURE 16 shows furthermore a centred bore hole (1) and two inclined opposite bore holes (2) which are leading out of the new shaft. Additionally there are two more horizontal bore holes (3) shifted at 90\u25e6 to the inclined bore hole (2) which are leading from the right to the left end. In order to find a possibility to get the thermocouple wires out of the steam environment, all these bore holes are required. The high live steam parameters and the high circumferential velocities of the rotor during the operation lead to high thermal and centrifugal loads on all components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002051_j.procir.2016.02.003-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002051_j.procir.2016.02.003-Figure4-1.png", + "caption": "Fig. 4: Process for the evaluation of the bearing clearance : a) Initial positioning of the bearing components b) Registration of the rollers on the outer ring c) Registration of the inner ring on the roller(s) in selected direction d) Registration of the inner ring on the roller(s) in the opposing direction", + "texts": [ + " For a good mesh quality it is necessary that the number of vertices respectively the length of the edges is sufficiently fine to represent the geometrical deviations. As especially waviness seems to be a severe problem for bearings (cf. section 3), the non-ideal geometries of the bearing components are considered to possess waves. Beside these systematic roundness deviations also random deviations (noise) and, of course, dimensional deviations can occur. Fig. 3 shows an example for a non-ideal geometry. The evaluation of the bearing clearance requires further sub-processes. In Fig. 4 the most important steps are presented. First of all a fixed center point is assumed serving as a reference for the following steps. The virtual non-ideal bearing components are then positioned around this center point according to the structure of the analyzed roller bearing. For this purpose each of the bearing components has its own center point defining its initial position. By contrast, the rotational orientation of bearing components is chosen randomly. As the bearing cage is neglected, the rollers are register on the raceway of the outer ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002574_detc2016-60019-Figure11-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002574_detc2016-60019-Figure11-1.png", + "caption": "Fig. 11 Superposition of CAD model on its virtual AM model", + "texts": [], + "surrounding_texts": [ + "For the Data transfer in coupled operations, MATLAB 14 programming language is used as a source system to write the contour lists file. The contour list data with correct data structure should be constructed for the successful coupling along with the CAD environment. Here, the target system is SOLIDWORKS macros. Therefore, contour list data will be written in a macros files format (.SWP files) as required in SOLIDWORKS. The process of writing of .SWP file data structure for the contour points and the corresponding height of extrusion or layer thickness includes following steps. Firstly, the contour lists are generated. Then, the vertices of each contour and corresponding constant extrusion height are written according to the .SWP file structure as shown in Fig. 7. Figure 7 displays the data format of .SWP file based on sliced contours lists. Set Part syntax opens a new part document in SOLIDWORKS software; Set skSegment syntax is used for modelling the contours and Set myFeature syntax is used for extrusion-purpose. e.g., if a first 2D contour has n points (i.e., P1, P2, P3, P4,\u2026Pi\u2026,Pn) the second contour has m points (i.e., P1, P2, P3, P4,\u2026Pj\u2026,Pm) at two different heights then according to .SWP file structure these should be written as Set skSegment = Part.SketchManager.CreateLine(P1, P2, P3, P4,\u2026Pi\u2026,Pn) Set skSegment = Part.SketchManager.CreateLine(P1, P2, P3, P4,\u2026Pj\u2026,Pm) Lastly, the complete file structure of 3D layered model from 2D contour lists in the form of .SWP file format is generated through MATLAB program and is used in the data transfer in coupled operations from the source system (i.e. MATLAB R2014a) to target system (i.e. SOLIDWORKS 2012)." + ] + }, + { + "image_filename": "designv11_34_0000781_mmar.2013.6669905-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000781_mmar.2013.6669905-Figure3-1.png", + "caption": "Fig. 3. Two Rotor Aero-dynamical System - parts scheme. Source: [5]", + "texts": [ + " The laboratory setup consists of the mechanical unit with power supply and interface to a PC and the dedicated RTDAC/USB2 I/O board configured in the Xilinx technology. The software operates in real time under MS Windows XP/7 32-bit using MATLAB R2009/10,11, Simulink and RTW toolboxes. Real-time is supported by the RT-CON toolbox from INTECO. Control experiments are programmed and executed in real-time in the MATLAB/Simulink environment. The real-life installation is presented in Fig.2, and the scheme of the system is presented on Fig.3. According to the TRAS instruction manual the equations describing the motion of the system can be written as follows: d\u2126v dt = lmFv(\u03c9m)\u2212 \u2126vkv + Uhkhv \u2212 a1\u2126vabs(\u03c9v) Jv \u00b7 \u00b7 \u00b7 + g((A\u2212B)cos\u03b1v \u2212 Csin\u03b1v) Jv \u00b7 \u00b7 \u00b7 \u2212 1 2 \u21262 h(A+B + C)sin2\u03b1vUhkhv Jv (3) d\u03b1v dt = \u2126v (4) dKh dt = Mh Jh = ltFh(\u03c9t)cos\u03b1v \u2212 \u2126hkh + Uvkvh Dsin2\u03b1v + Ecos2\u03b1v + F \u00b7 \u00b7 \u00b7 \u2212 a2\u2126habs(\u03c9h) Dsin2\u03b1v + Ecos2\u03b1v + F (5) d\u03b1h dt = \u2126h, \u2126h = Kh Jh(\u03b1v) , (6) and two equations describing the motion of rotors: Ih d\u03c9h dt = Uh \u2212H\u22121 h (\u03c9h) (7) and Iv d\u03c9v dt = Uv \u2212H\u22121 v (\u03c9v) (8) where: \u2126v - angular velocity (pitch velocity) of TRAS beam [rad/s]; \u2126h - angular velocity (azimuth velocity) of TRAS beam [rad/s]; \u03c9v - rotational speed of main rotor [rad/s]; \u03c9h - rotational speed of tail rotor [rad/s] Kh - horizontal angular momentum [N m s]; Mh - horizontal turning torque [ Nm]; Ih - moment of inertia of the main rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001080_s00170-013-4779-2-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001080_s00170-013-4779-2-Figure6-1.png", + "caption": "Fig. 6 PA-10 qi for i = 1, 2, \u00b7 \u00b7 \u00b7 , 7", + "texts": [ + " 24, we obtain V (k) = \u2212\u03b7(k)G2 N(k)||\u03c6(k)||2e2(k) 1 + \u03b7(k)G2 N(k)||\u03c6(k)||2 \u00d7 [ 1 \u2212 \u03b7(k)G2 N(k)||\u03c6(k)||2 2 + 2\u03b7(k)G2 N(k)||\u03c6(k)||2 ] . (32) With the learning rate in Eq. 21, the change of Lyapunov function candidate can be rearranged by V (k) = \u2212\u03b3 [ GN (k) Gu ]2 1 + \u03b3 [ GN (k) Gu ]2 \u23a1 \u23a2 \u23a2 \u23a3 1 \u2212 \u03b3 [ GN (k) Gu ]2 2 ( 1 + \u03b3 [ GN (k) Gu ]2 ) \u23a4 \u23a5 \u23a5 \u23a6 e2(k) \u2264 0. (33) Remark According to Eq. 33, the sign or direction of GN(k) is not needed. In this experimental setup, the proposed control algorithm is implemented to control a 7-DOF Mitsubishi PA-10 robotic arm system depicted in Fig. 5. With seven moving joints, Fig. 6 represents the positions for every joint variables q1, q2, \u00b7 \u00b7 \u00b7 , q7 used in this work. The overall system configuration is illustrated in Fig. 7. Seven FRENs are designed to control each of the joints independently. This robotic system is operated in the velocity-mode control which means that FRENs must generate the velocity commands to move every joint to follow the desired trajectory. To begin the controller design, the suitable IF\u2013THEN rules are all needed to be specified first. Based on the knowledge of PA-10, those IF\u2013THEN rules can be given as follows: If e(k) is PL Then u1(k) = \u03b2PL(k)\u03c61(k), If e(k) is PS Then u2(k) = \u03b2PS(k)\u03c62(k), If e(k) is Z Then u3(k) = \u03b2Z(k)\u03c63(k), If e(k) is NS Then u4(k) = \u03b2NS(k)\u03c64(k), If e(k) is NL Then u5(k) = \u03b2NL(k)\u03c65(k), where u denotes the velocity command determined by FREN and e(k) is the position error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003847_0954405420978120-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003847_0954405420978120-Figure2-1.png", + "caption": "Figure 2. 4D arc path: (a) minor arc and (b) major arc.", + "texts": [ + " D3 Pi,Pi 1\u00f0 \u00de= Pi, 3\u00bd Pi 1, 3\u00bd =ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xi xi 1\u00f0 \u00de2 + yi yi 1\u00f0 \u00de2 + zi zi 1\u00f0 \u00de2 q \u00f01\u00de Du Pi,Pi 1\u00f0 \u00de= Pi, u\u00bd Pi 1, u\u00bd = ui ui 1j j \u00f02\u00de Based on the definition of 4D representation and distance, a 4D arc path for SCARA robot path will be defined in Section 2.2. A circular arc in R3 can be defined by the arc center, start point, end point and arc normal. Extend this definition to 4D space: suppose O is the center point of a 4D arc C u\u00f0 \u00de, Ps is the start point, Pe is the end point, and n is the arc normal in R3. The 4D arc C u\u00f0 \u00de can be represented by equation (3). C u\u00f0 \u00de=O + sin ( 1 u\u00f0 \u00dea) OPs\u00f0 \u00de+ sin ua\u00f0 \u00de OPe\u00f0 \u00de sina , u 2 0, 1\u00bd \u00f03\u00de In equation (3), a is the rotation angle from Ps to Pe. As shown in Figure 2, the rotation from Ps to Pe may follow minor arc or major arc. Assume the angle between 3D vector OPs\u00f0 \u00de 3\u00bd and OPe\u00f0 \u00de 3\u00bd is u. The rotation angle a in equation (3) can be calculated by equation (4). a=u, if OPs\u00f0 \u00de 3\u00bd 3 OPe\u00f0 \u00de 3\u00bd : n. 0 a=2p u, else ( \u00f04\u00de Using the unified parametrization, the position portion of 4D circular arc is a 3D arc, and the rotation portion is a continuous quadratic curve of rotation angles relative to rotation arc length. Local arc transition and convex combination formulation Circular arc in R3 is often used for corner smoothing, because its arc length and curvature can be analytically computed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000342_gt2015-44069-Figure7-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000342_gt2015-44069-Figure7-1.png", + "caption": "Figure 7: Isometric view of the RTR", + "texts": [ + " Since the contribution of side bristles increases with increasing bristle deflection, the slope of the BTF-\u0394R curve increases continuously. However, the results of the FSS-TR measurements likely produce higher BTF values than actual turbine applications, where side effects are minimal due to large interference surfaces. 3 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/31/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use The high speed rotary test rig details are given in Figure 7. The rig mainly consists of seal housing assembly, spindle holder assembly and a solid base plate, also known as the \u201cplatform\u201d. The rig is powered by a GMN-high frequency spindle with 45,000 rpm maximum speed and 25 kW output rating. The torque is transmitted to the rotor with a connecting rod. The seal housing assembly is designed to test two seals at the same time (Figure 8). The housing assembly is positioned on the piezoelectric load cells such that lateral loading due to stiffness of brush seal can be measured during tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.12-1.png", + "caption": "Figure 9.12 A Stanford manipulator in its side and top views.", + "texts": [], + "surrounding_texts": [ + "A Stanford manipulator is shown in Figures 9.12 and 9.13. Its generic name originates from the research manipulator developed by Victor Scheinman at Stanford University. It comprises one prismatic joint and five revolute joints. Indicating the order of the joints as well, the manipulator is designated symbolically as 2R-P-3R or R2PR3. The joint axes are also shown in the figures together with the relevant unit vectors. The prismatic joint 3 and the succeeding revolute joint 4 form a cylindrical joint arrangement. Moreover, the last three revolute joints (4, 5, 6) form a spherical joint arrangement. The kinematic details (the joint variables and the constant geometric parameters) of the manipulator are shown in the line diagrams in Figure 9.13. The line diagrams comprise the side view, the top view, and two auxiliary views that show the joint variables that are not seen in the side and top views. The significant points of the manipulator are named as follows: O: Center point or neck point (origin of the base frame) S: Shoulder point, R: Wrist point, P: Tip point" + ] + }, + { + "image_filename": "designv11_34_0003557_elektro49696.2020.9130322-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003557_elektro49696.2020.9130322-Figure2-1.png", + "caption": "Fig. 2 Comparison between FEA and analytical approach. a) Magnetic flux density in the middle of motor b) Bn on line in the middle of air gap", + "texts": [ + " = 2 \u00b1 1, \u2208 \u2124 (10) \u2022 There are \u201cstep\u201d harmonic orders (11) \u2013 stator and (12) \u2013 rotor present in the spectrum, which are caused by the slotting of the rotor and stator. = = \u00b1 1, \u2208 \u2124 (11) = \u00b1 1, \u2208 \u2124 (12) \u2022 The winding factor for straight rotor bars is still given by equation (13) = 2 \u00b4 \u00b4 (13) Authorized licensed use limited to: East Carolina University. Downloaded on June 20,2021 at 07:38:56 UTC from IEEE Xplore. Restrictions apply. \u2022 According to formula (14), the magnitude of the harmonic can be calculated. = (14) These rules can be used for any distributed multiphase winding with integer . A FEM model was created for this motor to validate the method. Fig. 2 shows the geometry of the machine and a graph shows the distribution of the air gap flux density for the no-load state. The blue curve shows the analytical calculation, the red curve shows the result achieved by the FEA method. In order to compare both methods, the magnetic core is composed of an ideal steel with high relative permeability ( = 10 ). Fig. 2 shows that there is almost 100% agreement between these methods. When the induction machine is under load, voltage is induced into the rotor and a rotor current is generated. The magnetic field of this current interacts with the stator field and deforms the resulting mmf curve. FEM analysis was performed for the machine at its rated power = 3 . For the most accurate results, the magnetic core was composed of real steel with = ( ) . The results of this simulation are shown in Fig. 3. Due to iron saturation, spatial harmonics [9,10] arise in the motor, which are not predicted by rules (10-13) and are called \u201csaturation\u201d harmonics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001843_jahs.59.022006-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001843_jahs.59.022006-Figure8-1.png", + "caption": "Fig. 8. Whirl rig installed in vacuum chamber.", + "texts": [ + " This gives as close to a reading of internal PAM pressure as possible without a transducer inside of the actual PAM. To record the active and passive forces generated by the PAMs, a Honeywell Sensotec model 31 4500 N load cell was installed at the outboard end of each PAM, between the end fitting and the bell crank. Finally, to monitor the output flap deflection angle, a Midori noncontact Hall-effect sensor was installed at the outboard end of the flap. The body of the sensor was held fixed, whereas the shaft was coupled to the flap. This is shown in Fig. 8(a), which shows the outboard end of the test rig as installed in the whirl chamber. To maintain large deflection angles at higher frequencies, the design of the pneumatic supply system is very important for PAM TEF systems, as discussed in Ref. 36. Given the antagonistic configuration of this system, bidirectional operation of the flap system requires two PAM fill and exhaust cycles per flap cycle, one for each PAM as it is actuated in turn. Therefore, the pneumatic supply system effectively has to fill and exhaust the volume of one PAM and the length of tubing between that PAM and the control element at a rate of 70 Hz for 35 Hz flap actuation", + " Reducing tube diameter also reduces fill volume, but it has a negative effect on the obtainable flow rate. Studies on our system showed that 0.95-cm-outer-diameter (0.69-cm-inner-diameter) tubing provided the best balance of fill volume and flow rate. To supply air to the actuators, a rotating-frame air supply system was installed in the chamber. Because of the restricted size of the main rotor shaft on the rig and the presence of 60 signal and power wires running up through its hollow center, the airline was run up the wall of the chamber and then out to a mast on the whirl rig, as seen in Fig. 8(b). At the top of this mast was installed a commercially available pneumatic rotary union. This device is a low-cost, bearing-mounted, pressure-sealed rotating coupler that can transfer air from a fixed frame to a rotating frame at high flow rates and rotation rates up to 1500 rpm with low frictional losses. It is a well established and proven technology developed for demanding industrial applications. A house air compressor in the fixed frame (outside the chamber) then provided all the air required for testing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002961_0036850419897221-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002961_0036850419897221-Figure4-1.png", + "caption": "Figure 4. Brush seal model in Workbench. (a) Calculation process and (b) calculation model.", + "texts": [ + " In this article, because the calculated pressure is relatively small, the amount of deformation of the bristle is small, and the back baffle protection height is sufficient. In this case, it is considered that there is no significant influence on the deformation of the filament. Then, the solid and fluid models are separately suppressed during the process of meshing. In the fluid\u2013solid coupling calculation, the main calculation module used is system coupling that transfers the data between solid model and fluid model (Figure 4).24 To ensure the information transfer between solid and fluid models on the coupling surface without generating negative volume grids in the mesh deformation, tetrahedral cells are used in both solid and fluid domains. In the finite element (FE) calculation, nitrogen is used as the fluid medium and the RNG k\u2013e turbulence model is adopted, then the boundary conditions are set, which includes the pressure and temperature of inlet and outlet, the fluid solid coupling surface and dynamic mesh, the step of time, and the number of time steps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001343_pedstc.2013.6506688-FigureI-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001343_pedstc.2013.6506688-FigureI-1.png", + "caption": "Figure I. Model ofDMP machine", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION Every electrical machine whether it is AC or DC is a means of energy conversion. A common type of machine in the industry is a machine with one electrical port and one mechanical port. Their perfonnance is based on injecting electrical energy from their electrical port (usually stator) and obtaining mechanical energy from their mechanical port (rotor) in motor mode operation. DMP is a new generation of electrical machine consists of a stator and two rotors. This machine has electrical ports on the stator and inner rotor and mechanical ports on both rotors. Existence of two ports made these machines more flexible and increased their performance relative to previous ones. Because of their flexibility these machines could be applied in systems such as hybrid electric vehicles [1-6], wind energy conversion systems [7] and so on. In this paper, the proposed machine has two mechanical ports and two electrical ports. The paper is organized as follows: in the second part of this paper, structure ofDMP is probed. In the third part, the dynamic model and different modes of operation of DMP is studied. Simulation results are presented in part four. Finally in the last part conclusion is presented. II. STRUCTURE OF DMP DMP is more similar in structure to a DFIG with an extra outer rotor which makes it more flexible and controllable. In fact it is a compact machine that plays the role of two machines- inner machine consists of inner rotor and outer rotor and outer machine consists of outer rotor and stator. These two machines can function as either motor or generator. Depending on motor or generator function, different modes of operation are possible for DMP. Fig. 1 shows a typical DMP structure [8]. 978-1-4673-4484-5/13/$31.00 \u00a920 13 IEEE 125 Angular speed in one of the three parts is equal to zero and considered as stator. The two other rotational parts are considered as rotors. In this paper it is supposed that Wi is equal to zero and considered as stator. Part 2 and part 3 are outer rotor and inner rotor respectively. Fig. 2 shows these three discrete parts. The inner rotor is fed with a three-phase winding through slip rings. The stator is also fed with a three phase winding. Permanent magnets are mounted on the outer rotor. A rotating field is caused in the machine due to the existence of three-phase winding on the stator. As the outer rotor is a PM, it rotates synchronously with the stator rotating field like a PMSM operation. Inner rotor voltage frequency is also equal to the slip speed between the outer rotor and the inner rotor and another rotating field is caused in the machine due to the inner rotor. So the stator and the inner rotor fields rotate in the machine with the same speed. If Wz > UJ3 or UJz < UJ3, in both cases there will be a slip between the rotors. Energy flow in the inner rotor is proportional to slip of two rotors. Depending on the slip sign between the two rotors, direction of power flow between the two rotors will be positive or negative [8]. The rotors can be controlled by the control of stator electrical port. Moreover the machine can be controlled by the effect of external mechanical energy through mechanical ports (two rotor shafts). In the hybrid electric vehicle application, outer rotor of DMP is mechanically connected to the wheels and inner rotor is connected to the internal combustion engine (ICE). Electrical ports also have a shared DC link (battery). DMP machine is made in three forms namely radial flux, axial flux and combination of two previous cases. Different structures ofDMP machine are illustrated in fig. 3 [9]. III. OPERATION MODES OF DMP Energy balance equation in the DMP machine regardless of its losses is written as follows [1][3][8]: \u00b1Pe,stator \u00b1 Pm,pm-rotor \u00b1 Pe,wd-rotor \u00b1 Pm,wd-rotot = 0 (1) Where m indicates the mechanical energy and e indicates the electrical energy. Depending on the negative or positive sign, this equation describes different performances of DMP machine. In comparison with a common machine with one rotor, DMP has more operation modes. +Pe,stator - Pm,pm-rotor \u00b1 Pe,wd-rotor - Pm,wd-rotot = 0 (2) In equation (2), electrical energy is applied through stator ports and mechanical energy is received from mechanical ports of rotors. Depending on the relative speed between the rotors (outer rotor considered as reference) slip can be positive or negative, thus the flow of electrical energy to the electrical ports of the rotor can be positive or negative. In this condition DMP acts as a power splitter, because electrical energy of stator divides in various ports of the machine. -Pe,stator + Pm,pm-rotor \u00b1 Pe,wd-rotor + Pm,wd-rotot = 0 (3) In equation (3) unlike equation (2) mechanical energy of inner rotor considered as input of the machine, also transferred energy from the stator to the battery considered as the output of the machine. The direction of inner rotor electrical power could be positive or negative, depending on the slip sign of two rotors. In this condition DMP act as a power combiner. Because the mechanical energy of mechanical ports is combined together and appears as electrical energy in the electrical port of stator and can be saved in battery. This situation can be occurred in hybrid vehicles during deceleration. Consequently the mechanical energy of the vehicle and ICE will be saved in the battery. +Pe,stator - Pm,pm-rotor \u00b1 Pe,wd-rotor + Pm,wd-rotot = 0 (4) In equation (4), some constraints are applied to equation (1). In this condition mechanical energy and electrical energy are in equilibrium, and DMP shows performance of a gearbox with different gear ratios. The rotors are in energy equilibrium while speeds of rotors are different. Obviously, torque of the rotors are proportional too. In this situation, electrical energy of the inner rotor and the stator are in equilibrium and DMP acts as an EVT (Electrical Variable Transmission), because performance of CVT (Continuous Variable Transmission) is obtained by using an electromechanical appliance. Dynamic equations of DMP machine according to outer rotor reference frame are as follows: Aqs = Lsiqs + Lmiqr Ads = Am + Lsids + Lmidr Aqr = Lriqr + Lmiqs Adr = Am + Lridr + Lmids .ft (iqsAml + iqrAmz ) + Tout = P \ufffd (LSd - Lsq )idsiqs + \ufffd (Lrd - Lrq )idriqr + \ufffd (Lmd - Lmq )(idSiqr + iqsidr) (5) (6) (7) (8) (9) (10) (11) (12) (13) (14) Eqs. (1) through (4) are the voltage equations for the stator and the inner rotor windings. rs and rr are the stator winding resistance and the inner rotor winding resistance respectively. Eqs (5) through (8) are the flux linkage equations for stator and rotor windings. Ls and Lr are the self-inductance of the stator and rotor windings respectively. Lm is the mutual inductance between the inner rotor winding and the stator winding. Ami and Amz are the flux linkages produced by the PM-rotor and link the stator and inner rotor respectively. p is the number of pole pairs. Tout and Tin are the outer rotor and inner rotor torque respectively. According to the above equations, inner and outer machine are not independent but have interaction with each other. IV. SIMULATION In order to analyze the DMP dynamics a DMP with the following characteristics is simulated in an open loop test and without any controller. This machine has two pole pairs. TABLE!. PARAMETERS OF DMP[IO][II]" + ] + }, + { + "image_filename": "designv11_34_0002285_978-3-319-41009-8_59-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002285_978-3-319-41009-8_59-Figure2-1.png", + "caption": "Fig. 2. (a) Calculate the gradient direction when all robots have different fitness values. (b) Robots determine their own roles in role switching process. (b) Robots maintain the triangle formation with the aid of their internal compasses.", + "texts": [ + " \u2013 Case I: all three members share the same fitness value, which means that the team is in the area without fitness, and the leader will search randomly. \u2013 Case II: two robots share the same better fitness value, then the gradient vector equals the center of the two better positions minus the worse one. \u2013 Case III: two robots share the same worse fitness value, then the gradient vector equals the better position minus the center of the two worse ones. \u2013 Case IV: all robots have different fitness values, then the gradient is perpendicular to the local contour line (Fig. 2(a)). In Fig. 2(a), A, B and C stand for the robots in a team, and the fitness values satisfy the inequality f(A) > f(B) > f(C). Based on the local linear variation, f(B\u2032) = f(B) and the position of B\u2032 can be calculated from Eq. 1. The line BB\u2032 serves as a contour line, whose vertical vector B\u2032P is the gradient direction (Eq. 2). BB\u2032 = BC + CA \u00b7 f(B) \u2212 f(C) f(A) \u2212 f(C) . (1) { B\u2032P \u00b7 BB\u2032 = 0 B\u2032P \u00b7 B\u2032A > 0 . (2) Role Switching. In order to avoid turning abruptly, to facilitate the maintenance of formation and the estimation of gradient direction, a role switching trick is introduced. At each iteration, the robot with the maximum fitness value serves as the leader while the other two determine their roles (i.e. left wing or right wing) according to their relative positions. As is shown in Fig. 2(b), robot A is the leader and A\u2032 is its next position. AP is the right vertical vector of AA\u2032. If AB \u00b7 AP > AC \u00b7 AP , then robot B serves as the right wing, else the left wing. Formation Control. Each robot is assumed to be equipped with a compass. The leader will broadcast its next position within the team per iteration, and its members will determine their roles and next positions. As is shown in Fig. 2, given the positions of A and A\u2032, the positions of left wing and right wing (i.e. L and R) can be calculated. To maintain the formation, the leader will monitor the distances(D) from itself to its members, and slow down if the distance exceeds a certain threshold (T ), otherwise accelerate for high efficiency (Eq. 3, where \u03b1 and \u03b2 are factors for deceleration and acceleration). Parameters \u03b1 and \u03b2 are set to 0.75 and 1.33 respectively while T is set to 0.8 \u2217 Lengh, where Length (0.8 \u2217 2rt) is the ideal side length of the triangle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003066_j.promfg.2020.04.133-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003066_j.promfg.2020.04.133-Figure1-1.png", + "caption": "Fig. 1. FE-Model for RBPD", + "texts": [], + "surrounding_texts": [ + "ScienceDirect\nAvailable online at www.sciencedirect.com\nProcedia Manufacturing 47 (2020) 1134\u20131140\n2351-9789 \u00a9 2020 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 23rd International Conference on Material Forming. 10.1016/j.promfg.2020.04.133\n10.1016/j.promfg.2020.04.133 2351-9789\n\u00a9 2020 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (https://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 23rd International Conference on Material Forming.\nAvailable online at www.sciencedirect.com\nScienceDire t Procedia Manufacturing 00 (2019) 000\u2013000\nwww.elsevier.com/locate/procedia\n2351-9789 \u00a9 2020 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license https://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 23rd International Conference on Material Forming.\n23rd International Conference on Material Forming (ESAFORM 2020)\nRule-Based Path Identification for Direct Energy Deposition Moein Pakdel Sefidi a*, Rameez Israr a, Johannes Buhl a, Markus Bambach a\naChair of Mechanical Design and Manufacturing, Brandenburg University of Technology, Cottbus - Senftenberg, Konrad - Wachsmann - Allee 17, Cottbus, D-03046, Germany\n* Corresponding author. Tel.: +49355692913; fax: +49355693110. E-mail address: pakdelse@b-tu.de\nAbstract\nAdditive Manufacturing (AM) provides the possibility to produce complex part geometries on a layer-by-layer basis. In AM processes, a heat source moves over the part to either deposit powder or wire feedstock or to melt thin powder layers that are spread onto a powder bed. The moving heat source inevitably creates an inhomogeneous temperature distribution, which affects the residual stresses, part distortion, and local mechanical properties. Accumulation of heat in corners and path intersections may result in overheating and hence defects. In this study, a method for improving the temperature distribution in direct energy deposition processes is presented. FEM simulations in LS Dyna are coupled to MATLAB in order to divide a basic AM-FEM model and to adjust the model rule-based during the calculation. With this Rule-Based Path Identification (RBPI), the temperature history is used to choose the position of the next bead. With one bead wide wall, a temperature improvement of 90\u00b0C could be achieved. The new simulation method is adopted for a delay time between the beads for 10s. In conclusion, it is shown that the RBPD helps to reduce the inhomogeneous temperature distribution for a metal printing process without expensive optimization. \u00a9 2020 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license https://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 23rd International Conference on Material Forming.\nKeywords: Rule-Based Path Identification; RBPI; AM optimization; WAAM; DED.\nNomenclature\nAM Additive Manufacturing DED Direct Energy Deposition FEM Finite Element Method RBPI Rule-Based Path Identification SLM Selective Laser Melting WAAM Wire and Arc Additive Manufacturing\nDirect energy deposition, a subgroup of the additive manufacturing (AM) or 3D printing processes, comprises processes in which a part is made layer-by-layer by deposition of material [1,2]. The deposition process is enabled by a heat source that moves over the substrate on a predefined path and melts the feedstock material to produce the desired shapes [3,4]. DED processes are capable of producing complex three-\ndimensional parts with various geometrical features, which cannot be manufactured in conventional die-based processes [5,6]. The parts can be produced individually or in small batches [7]. Depending on the DED process the manufactured part may require subsequent machining [8]. In DED for metals, the material is added in the molten state and solidifies on the part. Due to heat transfer into the part, the temperature decreases rapidly, causing thermal shrinkage of the added material and the part which is being processed. Further, intensive localized heating increases the temperature gradient and leads to overheating in critical areas [14]. These effects result in high residual stresses and geometric distortion [9\u201311], as well as thermal cracks [12]. These unwanted phenomena depend on the spatial trajectory of the heat source as well as on the selected process parameters [12,13].\nWire- arc additive manufacturing WAAM is a DED process, which uses an electric arc as a heat source. The electrode is a wire that progressively consumes itself [15,16]. Molten\nAvailable online at www.sciencedirect.com\nScienceDirect Procedia Manufacturing 00 (2019) 000\u2013000\nwww.elsevier.com/locate/procedia\n2351-9789 \u00a9 2020 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license https://creativecommons.org/licenses/by-nc-nd/4.0/) Peer-review under responsibility of the scientific committee of the 23rd International Conference on Material Forming.\n23rd International Conference on Material Forming (ESAFORM 2020)\nRule-Based Path Identification for Direct Energy Deposition Moein Pakdel Sefidi a*, Rameez Israr a, Johannes Buhl a, Markus Bambach a\naChair of Mechanical Design and Manufacturing, Brandenburg University of Technology, Cottbus - Senftenberg, Konrad - Wachsmann - Allee 17, Cottbus, D-03046, Germany\n* Corresponding author. Tel.: +49355692913; fax: +49355693110. E-mail address: pakdelse@b-tu.de\nAbstract\nAdditive Manufacturing (AM) provides the possibility to produce complex part geometries on a layer-by-layer basis. In AM processes, a heat source moves over the part to either deposit powder or wire feedstock or to melt thin powder layers that are spread onto a powder bed. The moving heat source inevitably creates an inhomogeneous temperature distribution, which affects the residual stresses, part distortion, and local mechanical properties. Accumulation of heat in corners and path intersections may result in overheating and hence defects. In this study, a method for improving the temperature distribution in direct energy deposition processes is presented. FEM simulations in LS Dyna are coupled to MATLAB in order to divide a basic AM-FEM model and to adjust the model rule-based during the calculation. With this Rule-Based Path Identification (RBPI), the temperature history is used to choose the position of the next bead. With one bead wide wall, a temperature improvement of 90\u00b0C could be achieved. The new simulation method is adopted for a delay time between the beads for 10s. In conclusion, it is shown that the RBPD helps to reduce the inhomogeneous temperature distribution for a metal printing process without expensive optimization. \u00a9 2020 The Authors. Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND licens https://creativecommons.org/licenses/by- c-nd/4.0/) Peer-review under responsibility of the scientific committee of the 23rd International Conference on Material Forming.\nKeywords: Rule-Based Path Identification; RBPI; AM optimization; WAAM; DED.\nNomenclature\nAM Additive Manufacturing DED Direct Energy Deposition FEM Finite Element Method RBPI Rule-Based Path Identification SLM Selective Laser Melting WAAM Wire and Arc Additive Manufacturing\nDirect energy deposition, a subgroup of the additive manufacturing (AM) or 3D printing processes, comprises processes in which a part is made layer-by-layer by deposition of material [1,2]. The deposition process is enabled by a heat source that moves over the substrate on a predefined path and melts the feedstock material to produce the desired shapes [3,4]. DED processes are capable of producing complex three-\ndimensional parts with various geometrical features, which cannot be manufactured in conventional die-based processes [5,6]. The parts can be produced individually or in small batches [7]. Depending on the DED process the manufactured part may require subsequent machining [8]. In DED for metals, the material is added in the molten state and solidifies on the part. Due to heat transfer into the part, the temperature decreases rapidly, causing thermal shrinkage of the added material and the part which is being processed. Further, intensive localized heating increases the temperature gradient and leads to overheating in critical areas [14]. These effects result in high residual stresses and geometric distortion [9\u201311], as well as thermal cracks [12]. These unwanted phenomena depend on the spatial trajectory of the heat source as well as on the selected process parameters [12,13].\nWire- arc additive manufacturing WAAM is a DED process, which uses an electric arc as a heat source. The electrode is a wire that progressively consumes itself [15,16]. Molten", + "droplets are formed and transferred onto the solid substrate or the actual part. This causes a time-dependent temperature input into the shape. The droplets and the entire deposited weld bead undergo a remarkable temperature decrease. In comparison to the selective laser melting (SLM)process, the weld bead width is approximately 100 times larger in WAAM, i.e., about 180 \u00b5m in SLM [17] and 2-12 mm in WAAM [18] and the temperature distribution and resulting distortion are readily visible in WAAM after a few layers.\nIt is important to control the temperature during the WAAM process to obtain the desired shape and material properties. Previous work has shown the effect of different parameters with experiments and simulations and concluded that it would be highly beneficial to control the temperature during the welding processes. In this regard, FEM simulation is used to calculate the temperature distributions, distortions, strains and residual stresses [11,19,20].\nGibbons et al. [21] suggested that the process parameters \u2018welding path, welding parameters and type of heat source\u2018 should be properly chosen and pre-defined to achieve results with less geometrical inaccuracies. Using simulation, Zhang et al. [22] proposed that the local overheating during arc welding has a negative influence on the microstructure of the components. Hassel et al. [23] introduced pauses between the layers to avoid the global overheating of the component in their experiments. Using simulations, Rodrigues et al. [19] recommended substrate heating before welding to reduce the temperature gradient.\nIn general, the DED simulation is fully predefined, which means, that the process parameters are set before the simulation starts and cannot be adjusted during the simulation of the welding and the cooling processes. Hence, there is no possibility to control the parameters and prevent the unexpected rise in the temperature. Conventionally, numerical optimization with hundreds of simulation runs is used to obtain the optimized starting values of the process parameters [24]. However, a complex DED-simulation requires extremely high computational efforts and is prone to failure, which makes optimization highly expensive. Currently, it is impossible to control process parameters like the temperature of the AMmodel in a running FE simulation.\nThe aim of this study is to analyze the evolution of the temperature field after welding a single bead in order to define the trajectory of the next weld bead. Therefore, an observer setup with arbitrary rules is used to adjust the welding parameters and the path during the simulation. The observer inside the algorithm decides after a predefined increment if changes are needed and changes for example the direction of the weld path and establishes the new model for the next increment. Instead of a criterion-based termination, for example with the \u201cactuator-sensor interaction\u201d in Abaqus, this new way has the capability to interact (recognize, decide and choose the new position of the next bead). The aim of the paper is to explore rule-based simulation for identification of the welding path, which may be considered a step towards intelligent simulation.\nThe concept of RBPI is completely different from an optimization algorithm that requires a couple of simulations, which leads to extremely high computational costs for a fully defined DED-part. Using only a single run, the simulation with\nRBPI generates a welding path based on physical models with the process- and material knowledge that can be used as a start for experimental investigations.\nThe finite element model of the multi-layer wall on the substrate was built up with a basic path welding path. Ten layers of each five beads are stacked onto each other to build a wall (L= 60 mm, W= 4 mm, H= 18 mm). The substrate for deposition is a 2 mm thick metal plate of mild steel.\nThe welding path starts on the substrate from the right corner and moves in an alternating way with a constant velocity of 6.66mm/s without any stop until the 10th layer is welded. This path is time-optimized and well known in the industry.\nTo simulate the heat source for a WAAM-process, Goldak\u2019s double ellipsoidal model, based on Gaussian power density distribution, is used in this work [25]. The energy distribution in the forward, as well as the reverse direction can be given as:\n\ud835\udc5e\ud835\udc5e\ud835\udc53\ud835\udc53(\ud835\udc65\ud835\udc65, \ud835\udc66\ud835\udc66, \ud835\udc67\ud835\udc67) = 2\ud835\udc5b\ud835\udc5b\ud835\udc53\ud835\udc53\ud835\udc53\ud835\udc53\u221a\ud835\udc5b\ud835\udc5b\ud835\udc44\ud835\udc44 \ud835\udf0b\ud835\udf0b\u221a\ud835\udf0b\ud835\udf0b\ud835\udc4e\ud835\udc4e\ud835\udc53\ud835\udc53\ud835\udc4f\ud835\udc4f\ud835\udc53\ud835\udc53\ud835\udc50\ud835\udc50\ud835\udc53\ud835\udc53 \ud835\udc52\ud835\udc52\u2212\ud835\udc5b\ud835\udc5b(\ud835\udc65\ud835\udc65 \ud835\udc4e\ud835\udc4e) 2 \ud835\udc52\ud835\udc52\u2212\ud835\udc5b\ud835\udc5b(\ud835\udc66\ud835\udc66 \ud835\udc4f\ud835\udc4f) 2 \ud835\udc52\ud835\udc52\u2212\ud835\udc5b\ud835\udc5b( \ud835\udc67\ud835\udc67 \ud835\udc50\ud835\udc50\ud835\udc5f\ud835\udc5f )\n2\n(1)\n\ud835\udc5e\ud835\udc5e\ud835\udc5f\ud835\udc5f(\ud835\udc65\ud835\udc65, \ud835\udc66\ud835\udc66, \ud835\udc67\ud835\udc67) = 2\ud835\udc5b\ud835\udc5b\ud835\udc53\ud835\udc53\ud835\udc5f\ud835\udc5f\u221a\ud835\udc5b\ud835\udc5b\ud835\udc44\ud835\udc44 \ud835\udf0b\ud835\udf0b\u221a\ud835\udf0b\ud835\udf0b\ud835\udc4e\ud835\udc4e\ud835\udc5f\ud835\udc5f\ud835\udc4f\ud835\udc4f\ud835\udc5f\ud835\udc5f\ud835\udc50\ud835\udc50\ud835\udc5f\ud835\udc5f \ud835\udc52\ud835\udc52\u2212\ud835\udc5b\ud835\udc5b(\ud835\udc65\ud835\udc65 \ud835\udc4e\ud835\udc4e) 2 \ud835\udc52\ud835\udc52\u2212\ud835\udc5b\ud835\udc5b(\ud835\udc66\ud835\udc66 \ud835\udc4f\ud835\udc4f) 2 \ud835\udc52\ud835\udc52\u2212\ud835\udc5b\ud835\udc5b( \ud835\udc67\ud835\udc67 \ud835\udc50\ud835\udc50\ud835\udc5f\ud835\udc5f )\n2\n(2)\nWhere the weld pool energy can be expressed by:\n\ud835\udc44\ud835\udc44 = \ud835\udf02\ud835\udf02\ud835\udf02\ud835\udf02\ud835\udf02\ud835\udf02 (3)\nThe rest of the parameters of the above-mentioned equations have been adapted according to the previous work of Israr et al. [12] and are given in table 1.", + "1136 Moein Pakdel Sefidi et al. / Procedia Manufacturing 47 (2020) 1134\u20131140 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 3\nTo model the whole AM-part, every bead is considered as a separate part and a tied contact between the beads and the substrate is created at the start of the simulation. The thermal contact is established according to the welding sequence and depends on the adjacent beads.\nTherefore, if the welding sequence changes, the time of the contact activation changes, too. As the heat source passes over a specific bead, the thermal and mechanical properties of its elements are activated within the temperature range of 1400- 1450 \u00b0C. Details about the activation of the inactive thermal and mechanical material properties, as well as convergence analysis for coupled and uncoupled, pure and hybrid solvers, estimations of computation time, time step and mesh size can be found in Table 2 [26]. Following parameters are chosen for the simulation:\nTwo different material models from LS-DYNA have been used in this work. For the mechanical analysis of the beads, the generalized phase-change material *MAT_254 is used while for thermal analysis *MAT_Thermal_CWM T07 is used.\nIn contrast to optimization methods, which use many simulations to generate the optimized model, in this study the algorithm operates incrementally. A MATLAB script controls the simulation output. The workflow of the algorithm is shown in Fig. 3.\nThis first attempt of RBPI terminates the simulation at predefined sequences and extracts the results. After analysis and adjusting the process parameters, it generates new files to continue the simulation from the point of the previous termination. The idea of this paper is to identify the welding path step by step. In general, it is possible to repeat some increments if the deviation exceeds the target value, for example, if some areas in the AM part should follow a specific temperature.\nThe fully defined basic-AM model of a wall in LS Dyna is used as input for the RBPI. To identify the welding path, the basic AM model is structured into a main file which contains all unchangeable files like the material model and sequencedependent files like contact definitions.\nThe RBPI can be influenced by boundary conditions, i.e., sequences of specific beads and/or whole layers can be predefined. In the wall example, the first layer should be predefined by a consecutive bead sequence with their ID\u00b4s from" + ] + }, + { + "image_filename": "designv11_34_0000658_2015-01-0610-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000658_2015-01-0610-Figure1-1.png", + "caption": "Figure 1. The branched torsional vibration model of vehicle driveline", + "texts": [ + " Through the comparison of simulation, the impacts of gear meshing nonlinearities are analyzed. A 12 DOFs branched torsional vibration model of vehicle driveline considered the nonlinear frictional torque characteristics of clutch, dry frictional damping of friction plate, the meshing stiffness and backlash of two-stage gear pairs, differential, the different torsional stiffness of left and right half axle, the longitudinal adhesion characteristics of tire and some other factors, which is illustrated in Figure 1. In Figure 1, Te is engine torque, Tc is clutch frictional torque, Td is dry frictional damping torque of clutch friction plate. TLL and TLR are ground resistant torques of left and right driving wheels respectively. The frictional torque transmitted by the clutch in the launching process is given in Equation (1), (2), (3). The friction coefficient of clutch frictional material is influenced by their relative slipping velocity and the temperature of the frictional surface[9]. (1) (2) (3) Where \u03bcs is the coefficient of maximum static friction, v is relative slipping velocity between the friction surfaces, f(v) is the function of friction coefficient varied with relative slipping velocity, df(v)/dv is the slop of friction coefficient, T is the temperature of the friction surface, \u03b4(T) is the temperature influence function of friction coefficient, R is the equivalent frictional radius and Fn is the applied normal force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure12-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure12-1.png", + "caption": "Fig. 12. Fluid mesh outside the SEMMT-applied part.", + "texts": [ + " In the present study, the volume of the largest element in the SEMMT-applied part was calculated to be 6.2 \u00d7 10\u22124 and the volume of the smallest element outside the SEMMT-applied part was calculated to be 5.2 \u00d7 10\u22126. Therefore, C(e) is set to 1000 in the SEMMT-applied part, and is set to 1.0 outside the SEMMT-applied part. Fig. 11a, b, and 11c show the distribution of the element volume near the SEMMTapplied part. The red-dashed line in the figures indicates the boundary of the SEMMT-applied part. The outside of the SEMMT-applied part is divided into three bodies, as shown in Fig. 12. In each body, the unstructured elements are generated such that the element volume increases toward the outside to prevent the mesh distortion caused by small elements. This also decreases the number of elements and nodes, decreasing the computational cost. In this subsection, for the purpose of showing the effectiveness of the improved SEMMT, some analyses were conducted and the change of mesh quality was investigated. Two types of flapping motions, denoted by \u2018Motion A\u2019 and \u2018Motion B\u2019, are solved by three types of FSI analysis methods, denoted by \u2018Type 1\u2019, \u2018Type 2\u2019, and \u2018Type 3\u2019" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003295_s00170-020-05693-0-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003295_s00170-020-05693-0-Figure4-1.png", + "caption": "Fig. 4 Wheel path generation with a discretized method", + "texts": [ + " (7): T W h; \u03b8; t\u00f0 \u00de1 \u00bcT M G W h; \u03b8\u00f0 \u00de1 \u00bc R cos\u03b8 cos \u03c9 t\u00f0 \u00de\u2212R sin\u03b8 cos\u03b2t sin \u03c9 t\u00f0 \u00de \u00fe h sin\u03b2t sin \u03c9 t\u00f0 \u00de\u2212yt sin \u03c9 t\u00f0 \u00de R cos\u03b8 sin \u03c9 t\u00f0 \u00de \u00fe R sin\u03b8 cos\u03b2t cos \u03c9 t\u00f0 \u00de\u2212h sin\u03b2t cos \u03c9 t\u00f0 \u00de \u00fe yt cos \u03c9 t\u00f0 \u00de R sin\u03b8 sin\u03b2t \u00fe h cos\u03b2t \u00fe zt 1 2 64 3 75: 2 64 2 64 \u00f07\u00de It is noted that the relationship between the translation rate v and the angular velocity \u03c9 is directed by helical angle \u03bb can be expressed in Eq. (8): cot\u03bb \u00bc v rT \u03c9 : \u00f08\u00de For Eq. (7), the wheel\u2019s position and orientation [yt, zt, \u03b2t] are supposed to be determined to generate the wheel path. In the traditional conical flank grinding, the wheel path in the YT axis is approximated as a linear motion, which ignored the variation of wheel orientation and resulted in the grinding errors. In this work, to calculate the wheel position and orientation, the conical end-mill was discretized into a finite of segments shown in Fig. 4. Each step of grinding could be viewed as a cylindrical end-mill with a radius rTi. To this end, the wheel path for conical flank grinding could be simplified as solving a series of the flank grinding for cylindrical end-mill. For the cylindrical end-mills, to determine the wheel\u2019s position, a grinding point is introduced. The grinding point is located by the parameter HGPi shown in the broken view of Fig. 5. Based on the grinding point, the wheel\u2019s position [yi zi] can be represented in Eq. (9): yi \u00bc rTi\u2212HGPi sin\u03b2i \u00fe R cos\u03b2i zi \u00bc \u2212HGPi cos\u03b2i \u00fe R sin\u03b2i : \u00f09\u00de Besides, the cylindrical flank grinding can be expressed as Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003675_ecce44975.2020.9236229-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003675_ecce44975.2020.9236229-Figure1-1.png", + "caption": "Fig. 1: A pair of coils (solid thick lines) on two concentric baseline circles (dashed lines) of radii rb1 and rb2. The grey circles represent the cross-sections of the coils. The angle between two coil axes is \u0394\u03b8.", + "texts": [ + " Supposing the coil is composed of N straight segments, the self-inductance of the coil is Lcoil = N\u2211 i=1 \u239b \u239dLi + N\u2211 j=1,j =i Mij \u239e \u23a0 (3) where Li is the self-inductance of i-th segment, Mij is the mutual inductance between the i-th and the j-th (i = j) segments. Supposing the coil 1 and coil 2 are composed of N1 and N2 straight segments respectively. The mutual inductance between two coils is Mcoil = N1\u2211 i=1 N2\u2211 j=1 Mij (4) where Mij is the mutual inductance between the i-th segment from coil 1 and the j-th segment from coil 2. In a radial air-cored machine, stator and rotor coils are typically located on two concentric baseline circles of radii rb1 and rb2, as shown in Fig. 1. When the coil geometries and the radii of the baseline circles are fixed, the mutual position of the coils is defined by the angle \u0394\u03b8 between the two coil axes, which is the only variable influencing the mutual inductance. As a result, the mutual inductance of a pair of coils can be written as Mcoil = M(\u0394\u03b8) (5) which is a function of the angle between the two coils. 5814 Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 17,2021 at 22:47:49 UTC from IEEE Xplore. Restrictions apply. The next step is to group the individual coil inductances to form the phase winding self and mutual inductances. Balanced three-phase, single-layer, lap windings are assumed for the stator and rotor with each winding composed of identical coils with appropriate connections. Stator and rotor are different, being laying on concentric baseline circles of different radius (see Fig. 1). The winding layout has periodicity pp: this property can be used to reduce the number of calculations for coil-to-coil inductances. The self-inductance, Lph, of the stator or rotor phase (i.e. with identical coils on the same baseline circle) can be divided into three parts, Lph = L0 +M1 +M2 (6) where L0 is the sum of self-inductances of all coils in the phase, M1 is the sum of mutual inductances between coils in the same phase and same pole pair, M2 is the sum of mutual inductances between coils in the same phase but different pole pairs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001020_icems.2014.7013720-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001020_icems.2014.7013720-Figure3-1.png", + "caption": "Fig. 3. Location distribution of d axis and q axis when winding A and C is excited under star connection. (a) measuring position for d-axis inductance. (b) measuring position for q-axis inductance.", + "texts": [ + " From (4), it can be derived that 2 2 AA CC AC / 22 2 2 U I RL L M (5) Substituting (3) into (5), we can obtain the following equation: 1539 2 2 0 0 2 / 23 2\u03c0cos 2 2 3 2s s s U I R L M L (6) Comparing (1) with (6), the d- and q-axis inductance of PMSLM can also be expressed as 2 2 max 2 2 min 2 2 2 2 d q U I R L U I R L (7) where Umax is the maximum of the line voltage when \u03b8=\u03c0/6, Umin is the minimum of the line voltage when \u03b8=2\u03c0/3. Through the above analysis, the d- and q-axis inductance of PMSLM can be obtained only by measuring the basic electrical parameters at two special rotor positions when arbitrary two phase windings are energized by alternating current. Fig. 3 shows the two special positions for measuring the d- and q-axis inductance of PMSLM. Coincidentally, the line voltage reaches its extreme value at the two positions. Therefore, we only need to detect the line voltage instead of the accurate positioning in advance. To obtain reference values of inductance, a 3-D finite element model of PMLSM, shown in Fig. 4, is built in commercial FE software JMAG. The parameters are shown in Tab. I. Fig. 5 shows the simulation results of self-inductance and mutual inductance and the values are shown in Tab" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002875_s42496-020-00033-7-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002875_s42496-020-00033-7-Figure4-1.png", + "caption": "Fig. 4 MSC Adams dynamic model", + "texts": [ + " According to such an approach, the allowed relative rigid rotations between each section of the panels are reproduced as an increased flexibility of the overall array. (4)mi = mf [ tanh ( 2 ih / ) i ( 2 i \u2212 1 ) h ] , (5)hi = h 2 \u2212 2 i [ tanh ( ih ) \u2212 1 \u2212 cosh ( 2 ih / ) sinh ( 2 ih\u2215 ) ] , (6)ki = mig 2 i tanh ( 2 ih ) , (7) 1 = 1.841, 2 = 5.329, i \u2243 i\u22121 + . (8) I0 = ( 1 \u2212 0.85 h \ud835\udf11 ) mf ( 3\ud835\udf112 16 + h2 12 ) \u2212 m0h 2 0 \u2212 3 \u2211 i=1 mih 2 i , if h \ud835\udf11 < 1 1 3 The spacecraft is hereby represented as a parallelepiped platform with solar panels as appendages (respectively, in green and in red in Fig.\u00a04). The two tanks are modelled as bodies with assigned mass and inertia to reproduce a hollow cylindrical tank. The flexibility of the two solar panels is simulated using FEM techniques. In particular, the two arrays are shell bodies meshed with Quad4 elements. The two tanks are rigidly connected to the platform to simulate their fastening to the satellite. A detailed image of a single tank is provided in Fig.\u00a05. The sloshing masses ( m1 in green, m2 in blue and m3 in red) have been modelled as\u00a0spherical bodies with assigned mass deriving from Eq", + " Such masses have been connected via translational joints to the walls of the tank, so that they are constrained to move only in the horizontal direction to simulate the lateral sloshing behaviour. At the same time, two spring\u2013damper forces are acting on each mass to reproduce the elastic forces they are subjected when moving from their equilibrium position (these forces are representative of the \u201celastic\u201d properties of the fluid). The stiffness of each spring is computed using Eq.\u00a0(6). The red sphere in the lower part of the tank in Fig.\u00a04 represents m0 . Its position and value are derived from Eqs.\u00a0(1)\u2013(2), while its moment of inertia from Eq.\u00a0(8). In most Earth Observation missions, the spacecraft is required to sweep a specific area with the same orientation orbit after orbit. One possible objective is to use the scientific instrumentation to acquire data of a strip, parallel to the satellite ground track (see Fig.\u00a06). In this case, the mission can be divided into two phases: a first attitude manoeuvre 1 3 to re-orient the payload and a sensing phase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003390_1350650120945517-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003390_1350650120945517-Figure4-1.png", + "caption": "Figure 4. Equivalent linear dynamic model of lubricant film.", + "texts": [ + " R3 L FX FY \" # \u00bc XL i\u00bc1 Fi X Fi Y \" # \u00bc XL i\u00bc1 Z \u00fe1 1 Z i 2 i 1 pi cos sin d dz \u00f016\u00de As it stated in equation (17), dynamic forces are a combination of static and perturbed forces FX FY \u00bc FX0 FY0 \u00fe FX FY \u00f017\u00de In the linear dynamic analysis model, perturbed forces FX, FY\u00f0 \u00de are considered as a function of the dynamic components of rotor displacement X0,Y0\u00f0 \u00de and velocity _X0, _Y0 . In this way, the equivalent stiffness and damping coefficients of the lubricant film can be calculated for the assumed dynamic rotorbearing system as depicted in Figure 4 according to the following equation FX FY \u00bc SXX SXY SYX SYY X0 Y0 BXX BXY BYX BYY _X0 _Y 0 ( ) \u00f018\u00de where Sij,Bij i, j \u00bc x, y\u00f0 \u00de represent the equivalent stiffness and damping coefficients of the couple stress lubricant film, respectively. Further, the magnitude of these dynamic coefficients can be calculated as follows SXX SXY SYX SYY \u00bc 2 C3 m ! R3 L Sij \u00bc XL i\u00bc1 Z \u00fe1 1 Z i 2 i 1 px py cos sin d dz \u00f019\u00de BXX BXY BYX BYY \u00bc 2 C3 m R3 L Bij \u00bc XL i\u00bc1 Z \u00fe1 1 Z i 2 i 1 p _x p _y cos sin d dz \u00f020\u00de Linear stability margin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003489_icirca48905.2020.9183199-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003489_icirca48905.2020.9183199-Figure1-1.png", + "caption": "Fig. 1. Aircraft body-fixed coordinate system", + "texts": [ + " Various moments and forces comprising aerodynamic, propulsive and gravitational elements govern the equations of motion of aircraft. These forces and moments operate at the aircraft center of mass. Newton\u2019s Laws of Motion is employed to derive the equations of aircraft motion [2]. Orientation and position of the aircraft are represented with respect to body-fixed coordinate system [1]. The threedimensional system of coordinates is denoted by three mutually perpendicular axes x-y-z. The origin of coordinate system coincides with the center of mass of aircraft as depicted in Fig. 1. The three rotational motions yaw, pitch and roll are about these axes. Yaw is the rotational motion about vertical (z) axis, about longitudinal (x) axis is roll and pitch is the rotation about lateral (y) axis. Primary flight controls of aircraft are elevator, rudder and ailerons. Roll is controlled by ailerons in each wing on the outer rear edge. Elevator controls the pitch on the horizontal tail surface and yaw is controlled by rudder on the vertical tail fin. The aircraft equations of motion which are nonlinear can be decoupled into two sets designated as lateral- directional motion and longitudinal motion", + "00 \u00a92020 IEEE 428 Authorized licensed use limited to: University of Durham. Downloaded on September 27,2020 at 03:31:47 UTC from IEEE Xplore. Restrictions apply. Therefore, only longitudinal dynamics is studied for pitch control in this work. The four forces of flight are drag, lift, thrust and weight. These forces balance out each other under the assumption of steady cruise. The two orientation angles are angle of attack (\u03b1) and side slip angle (\u03b2). They are used to describe aerodynamic moments and forces, also recognized as aerodynamic angles . Fig. 1 illustrates the orientation angles, velocity elements, forces and moments in body-fixed axes. X, Y and Z represent aerodynamic force elements, L, M and N represent moment components and \u03c8, \u03b8, \u0424 and \u03b4\u2091 denote yaw, pitch, roll and elevator deflection angles [1]. The notations u, v and w denote velocity components, while p, q and r represent angular rates about roll, pitch and yaw axes respectively. Further to the assumption of steady flight, no speed change due to change in pitch angle is considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure8.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure8.2-1.png", + "caption": "Figure 8.2 A pencil-like end-effector.", + "texts": [ + " On the other hand, in a manipulator without a spherical wrist, the functions of the joints cannot be separated distinctly as explained above. In other words, the location of the wrist point with respect to the base frame is determined not only by the first m\u2212 3 joints but also by the antepenultimate joint m\u22122. Besides, in some special configurations such that the approach vector u\u20d7a of the end-effector is not aligned with the axis vector u\u20d7(m) 3 of the last joint m, even the penultimate joint m\u22121 becomes effective on the location of the wrist point R. Such a configuration is illustrated in Figure 8.2. Note that, in such a configuration, the wrist point R happens to be located at Om instead of the usual Om\u2212 1. The preceding statements about the spherical and nonspherical wrists can be expressed by the following equations that show the functional relationships between the joint variables and the orientation of the end-effector and the location of the wrist point with respect to the base frame. Position and Motion Analyses of Generic Serial Manipulators 207 For a manipulator with a general nonspherical wrist, r = f (q1, q2,\u2026 , qm\u22123, qm\u22122, qm\u22121) (8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002256_978-3-319-25178-3_4-Figure4.4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002256_978-3-319-25178-3_4-Figure4.4-1.png", + "caption": "Fig. 4.4 Passive tag load modulation (Cisco 2008)", + "texts": [ + " Low-cost tags, applicable to the grocery industry, cost from 20 cents to 35 cents, and the latest tag developments promise tags that will cost just around 5 cents (Ka\u0308rkka\u0308inen 2003). However, tags can also cost several dollars depending on many factors such as: \u2022 Data capacity \u2022 Form \u2022 Operating frequency \u2022 Range \u2022 Performance requirements \u2022 Presence or absence of a microchip \u2022 Read/write memory Passive RFID tags vary in how they broadcast to RFID readers and how they receive power from the RFID reader\u2019s inductive or electromagnetic field. This is commonly performed by two basic methods: \u2022 Load modulation and inductive coupling in the near field as shown in Fig. 4.4. The RFID reader provides a short-range alternating current magnetic field that the passive RFID tag uses for both power and broadcasting. Through inductive (near field) coupling, the magnetic field induces a voltage in the antenna coil of the RFID tag, which powers the tag. The tag broadcasts its information to the RFID reader. Each time the tag draws energy from the RFID reader\u2019s magnetic field, the RFID reader itself detects a corresponding voltage drop across its antenna leads. Thus, the tag can communicate binary information to the reader by switching a load resistor on and off to perform the load modulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003506_b978-0-12-817582-8.00022-2-Figure15.1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003506_b978-0-12-817582-8.00022-2-Figure15.1-1.png", + "caption": "FIGURE 15.1 Single-link flexible joint manipulator.", + "texts": [ + " The trajectory tracking results have been discussed in Section 15.5. Finally, conclusions are presented in Section 15.6. In the era of industrial automation, industrial robot manipulator is a major development that has helped in increasing the production rate and quality of products simultaneously. Robot motion control is a problem that has been analyzed for decades now in order to improve the robot performance, reduce the robot cost, improve safety, while introducing new functionalities. Thus, the tracking problem of a SLFJM has been considered. Fig. 15.1 represents the model of a typical single-link flexible joint manipulator. The system equations for this model are: Il \u03b8\u0308 + mgl sin \u03b8 + k (\u03b8 \u2212 \u03b8m) = 0 (15.1) Im\u03b8\u0308m \u2212 k (\u03b8 \u2212 \u03b8m) + \u03bc\u03b8\u0307m = u (15.2) where \u03b8 is the link angle, \u03b8m is the angle of rotation of the motor, Il is the inertia of link, Im is the inertia of motor actuator, k is the elastic stiffness of the flexible link, m is the mass of the link, g is the acceleration due to gravity, l is the link length, \u03bc is the viscosity, and u is the control torque generated by the motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002817_j.apm.2020.01.007-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002817_j.apm.2020.01.007-Figure1-1.png", + "caption": "Fig. 1. Planar adjoint approach for planar four-bar linkage.", + "texts": [ + " [31] used the image space to represent the spherical motion by kinematic mapping. The complicated equations and redundant variables of the algebraic methods bring difficulties into in-depth study of geometric properties. A unified geometric method is preferred to describe the spherical coupler curve, based on which the local and global properties can be both revealed. For any planar four-bar linkage, the hinge point B traces a circle or circular arc B , while the coupler point P traces the coupler curve in the frame link, as shown in Fig. 1 . Through Cesaro\u2019s adjoint approach in differential geometry P [32] , P can be viewed as the adjoint curve to the original curve B . The Frenet frame { r B ; e 1 , e 2 } of B is set up to describe the position vector of P . Since B is a simple circle, the vector equation of P and its derivatives can be concise and easy to implement. One of the authors [33] expanded the Cesaro\u2019s planar adjoint approach to general spherical motion and described the spherical kinematics in differential geometry language" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000788_metroaerospace.2014.6865903-Figure4-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000788_metroaerospace.2014.6865903-Figure4-1.png", + "caption": "Fig. 4. Case study: blueprint.", + "texts": [ + " Min-max range computes the distance between the face centroids on the overall amount of points and reports it together with minimum and maximum distances that are found adding the distances of the two couples of the most outer and inner points of the faces. Percentile analysis gives a deeper overview of the distance distribution since it computes the averaged distance as the distance between the median position (taken from the best fitting planes) of each face and the maximum and minimum values as percentile distances (10% or 90%). This allows to cut off outliers or couples of points that are not statistical relevant for the measurements. III. CASE STUDY The proposed application aims to apply the procedure to an aeronatical flange shown in Fig. 4 and Fig. 5. It is a component of a Boeing product, obtained through a \u03b2-forging process starting from Ti6Al4V Titanium Alloy Fig. 3. Example of dimensional and planarity inspection of two planar surface. powders, that is being studied by Centro Sviluppo Materiali. This powder material is suited for manufacturing near net shape forged components since it has been demonstrated that it is possible to manufacture by powder-metallurgy samples that are mechanically equivalent to those from commercial extruded bar [10, 11]", + " 6 shows that planar surfaces always have standard deviation smaller then the cilindric ones, with a ratio of about 1:2. This allows to find a significant threshold for distinguishing cylinders from planes. The IRS is found starting from the direction orthogonal to the transversal section of Fig. 6. The second axis is found as the most populated set of aggregated voxels with a direction normal to the first one. It results to be one of the set of the five protrusions at 72\u00b0, as shown in the blueprint of Fig.4. The third axis derives from vectorial product. From IRS the recognition of cylindric voxels is found through the osculating radius evaluation. Fig. 7 shows the map of the voxel\u2019s osculating radii of the sections with smaller holes, confirming a good resolution of the proposed method. For example considering the 4 smaller holes in the left part of Fig.7 the voxel evaluation found averaged values nearby 6 mm, while the 8 holes on the right are around 5 mm. A preliminar validation of the voxel\u2019s osculating radius algorithm has been made comparing the voxel evaluation with surface reconstructions made in CATIAV5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000382_icra.2015.7139501-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000382_icra.2015.7139501-Figure3-1.png", + "caption": "Fig. 3. Brachytherapy procedure simulation setup including ultrasound probe, phantom tissue, needle template, and needle. The setup is represented from the field of view of the camera.", + "texts": [ + " This technique for segmenting a curved needle in an ultrasound is then extended to a camera image. The needle segmentation algorithm results in a second order polynomial equation that follows the curvature of the needle in its respective frame of reference, i.e. either the camera frame or the ultrasound image frame. The experimental setup consists of a tissue phantom imaged from overhead by a camera and from the side by an ultrasound probe. The images show a needle that is inserted into the tissue that deflected during insertion. Fig. 3 shows a schematic representation of the experimental set up in order to define the different coordinate frames that arise from combining imaging modalities. Points in both the camera and ultrasound images are throughout the paper defined in their respective pixel domains. In the convention here, superscripts and subscripts on the left side of a symbol refer to the frame of reference for the coordinate system. The coordinate frame for the ultrasound image is denoted by U ; a point U P in the ultrasound image will have coordinates (U x, U y) defined with respect to the lower left hand side of the ultrasound image, as shown in Fig. 3. The coordinate frame for the camera image is denoted by C; a point CP in the camera image will have coordinates (Cx, Cy), defined with respect to the upper left hand side of the camera image, as indicated in Fig. 3. A third frame is defined not in a pixel domain but in physical coordinates and referred to as the real-world frame, denoted by R. A point RP in the real-world frame consists of (Rx, Ry) measured in meters with respect to an origin located where the needle exits the template (see Fig. 3). Needle segmentation in 2D ultrasound images has previously been done via Gabor Filtering [6] and the Hough Transform for straight needles [7] or curved needles [8]. Additionally, needles have been segmented in 3D ultrasound images using projection methods [9], [10] or RANSAC [11]. In this paper, we combine RANSAC filtering with Otsu\u2019s Algorithm [12]. Ultrasound images along the length of the needle are taken near the needle\u2019s entry point into tissue, as this represents the most challenging scenario for prediction", + " The matrix R U T that transforms points from the ultrasound frame to the real-world frame is of the form R U T = R U Sx cos\u03c6 sin\u03c6 R U tx \u2212sin\u03c6 R U Sy cos\u03c6 R U ty 0 0 1 (10) where R U tx and R U ty are the x and y offset, in meters, of the leftmost point of the transducer element from the base of the needle, \u03c6 is defined as the angle of the ultrasound probe relative to the closest edge of the tissue, and R U Sx is the pixel domain width scaling factor and R U Sy is the height scaling factor, see Fig. 3. The scaling elements of the matrix are constant in each tissue and can be found by measuring object distances within the tissue. The translation elements can be found through measuring the offset of the ultrasound probe. In contrast we will discuss how we calculate the value of \u03c6 in Section IV-C as this requires having the matrix R CT . The matrix R CT that transforms points from the camera frame to the real-world frame is given by R CT = R CS0 0 R Ctx 0 R CS0 R Cty 0 0 1 (11) where R Ctx and R Cty are the x and y offset, in meters, of the upper leftmost point of the camera image to the base of the needle and R CS0 is the pixel domain scaling factor for both the x and y axes as the experiment is centered in the camera\u2019s field of view" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002915_s10846-020-01168-2-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002915_s10846-020-01168-2-Figure1-1.png", + "caption": "Fig. 1 Two common types of handles for the robot walking helper: (a) horizontal and (b) straight", + "texts": [ + " Meanwhile, 12 participants are with more than 2 years of experience in using the robot walking helper, 4 with 8 to 24 months, 6 with 1 to 8 months, and 44 with no prior experience. We first summarize the main human factors that have been discussed in previous researches [6\u201314], including (1) psychological factors: visibility, distance to the obstacle (wall, crowd, etc.), walking habit (right or left side), and (2) physical factors: moving and turning speeds, jerk. Specific for the robot walking helper itself, we also include its outlook and the types of handles for the user to hold. Fig. 1 shows two common types of handles, i.e., horizontal and straight [16]. Because the users need to adjust their postures when holding each of these two types of handles to push the robot walking helper, it would solicit different degrees of comfort feeling, and thus deserves investigation. During the survey, the participants were asked to answer the following two questions: 1) What factors are most important to you for comfortbased motion in using a robot walking helper? Please list at least three" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002078_s1995421215040085-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002078_s1995421215040085-Figure1-1.png", + "caption": "Fig. 1. Velocity contours in a flow: (a) at stream course motion, (b) at countermotion of the stream course and liquid flow, (c) at one way stream course motion and liquid flow. \u03c4 and (\u03c4 + d\u03c4): shearing strain of the upper and lower edge of the fluid flow element.", + "texts": [ + " The quality of the polymeric coating depends on the microstructure, and it, in turn, is defined by the technology used [6, 7]. The aim of theoretical investigations is to deter mine the conditions for forming a uniform polymeric coating on the external surface of a rotating cylinder. It should be noted that formation of the polymeric coating is carried out in the first and fourth quadrants of the rotating cylinder surface, i.e., in areas with countermotion of the fluid flow and stream course. Let us analyze the velocity contour at countermotion of the fluid flow at speed V and the stream course at speed u (Fig. 1b). At point A, the flow rate is 0. Below point A, the fluid layer moves toward the stream course due to the friction free flow. Above point A, the fluid layer moves toward the fluid flow due to pressure difference Pf. The layer of the polymeric material solution at the point A level has the best condition for polymerization. The level (y) of point A relative to height of the flow a; i.e., the y/a ratio is of great importance. Figure 2 shows the velocity contours at the counter and trail motion of the fluid flow and stream course with the ratio of the level of point A relative to the flow height y/a = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000868_2013-01-2353-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000868_2013-01-2353-Figure1-1.png", + "caption": "Figure 1. A typical three-axle vehicle with tandem rear axles.", + "texts": [ + " For a general vehicle these equivalent wheelbase and understeer parameters have been derived [5,6], Equation 3 Equation 4 where \u03b7i is the constant of proportionality (or \u201cgain\u201d) between the displacement of the ith axle and the primary steered front axle. If the ith axle is not steered \u03b7i is zero. The xi in Equation 3 and Equation 4 are the signed locations of the n axles. An axle in front of the vehicle center of mass has postitive location, and if the axle is behind the center of mass it will have a negative location. In this work we focus on the three-axle vehicle in its most common configuration of a single primary steered axle at the front, and two closely coupled tandem axles at the rear, as shown in Figure 1. The fundamental forces that determine a vehicle's handling response are generated by the tires. A tire generates a lateral force in response to a slip angle, that is the difference between the direction the tire is pointed and the direction it is moving. As shown in Fig 2, the direction a given tire is moving is a function of the vehicle's longitudinal and lateral velocities, and of its yaw rate. The side force coefficient Ci found in Equation 3 and Equation 4 multiplies the slip angle of a tire to yield its lateral force. In the two-degree-of-freedom model all tires experience the same slip angles, so their action can be combined for an overall effect at each axle. By summing forces and moments and considerable mathematical manipulation Equation 2, Equation 3 and Equation 4 are derived [5,6]. Special cases of Equation 3 and Equation 4 can be written for the three-axle vehicle of Figure 1. Equation 5 Equation 6 where l is the overall wheelbase Equation 7 and t is the tandem spread between the rear axle. Equation 8 Modification for Normal Force The typical two-degree-of-freedom vehicle model assumes a proportional relationship between the slip angle of a tire and the lateral force it generates. For small slip angles and relatively constant vertical loads on the tire this is a good assumption. This assumption does not allow analysis of a caster-steered axle, however. The characterization of lateral force produced by a tire can be quite complicated [12] but for the purposes of this work a simple tire model can be used, Equation 9 where the side force coefficient Ci of each axle i is a function of the nominal side force coefficient corresponding to a nominal normal load No, and the actual normal load carried by that axle Ni" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-FigureD.14-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-FigureD.14-1.png", + "caption": "Fig. D.14 Cylinder as a particular case of an ellipsoidal prism", + "texts": [ + "13 (The inertia tensor of a ellipsoidal plate) An ellipsoidal plate is a particular case of an elliptical prism for which the height of the prism is very small compared to the ellipsoidal face dimensions. Then its inertia moment matrix arise after (D.24) assuming a \u2248 0 (Fig.D.13): Example D.14 (The cylinder) The cylinder is also a particular case of the elliptical prism where both main diagonal components of the elliptical plane are equal. For instance b = c = r implies an horizontally placed cylinder of radius r and length l = A = 2a, with volume V = \u03c0r2l, and diameter D = B = C = 2r . The inertia matrix is found after these conditions over expression (D.24) (Fig.D.14). 430 Appendix D: Examples for the Center of Mass and Inertia Tensors of Basic Shapes I c = m 12 \u23a1 \u23a3 6r2 0 0 0 l2 + 3r2 0 0 0 l2 + 3r2 \u23a4 \u23a6 = m 48 \u23a1 \u23a3 6D2 0 0 0 4l2 + 3D2 0 0 0 4l2 + 3D2 \u23a4 \u23a6 (D.25) Example D.15 (The inertia tensor of a pole) A pole is the particular case of a cylinder with a very small diameter (radius) in comparison with the length: r l. In this case the mass distribution in the radial direction is negligible and the Inertia moments matrix can be computed using (D.25) assuming r \u2248 0 (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000217_978-3-030-04275-2-Figure8.2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000217_978-3-030-04275-2-Figure8.2-1.png", + "caption": "Fig. 8.2 Sketch for a 3-wheeled omnidirectional robot where the contact force, the power torque and a virtual frame in the first wheel are shown", + "texts": [ + "70), while the gravity wrench, consider that the non-inertial frame is placed with the z-axis pointing upwards and the roll-pitch-yaw attitude representation (R-P-Y ) is chosen, as in Example6.2. Then, the gravity vector is g0 = \u239b \u239d 0 0 \u2212g \u239e \u23a0 (8.17) and the gravity wrench (\u2212MG(1)) is expressed by (6.103). 336 8 Model Reduction Under Motion Constraint The Exogenous Forces F Consider that this robot has three omnidirectional wheels, like in Fig. 8.3-right. For simplification purposes, consider the wheels to be mass-less bodies and the contact point between the floor and the wheel to be constant relative to the wheel position (refer to Fig. 8.2). For kinematic analysis, consider the sketch of Fig. 8.4, where the non-inertial frame is placed at the geometric center of the wheels such that the x-axis is positive to the front of the vehicle, the z-axis is positive in the upward direction and the y-axis is placed according to the right-hand rule. Consider three local frames ri at the contact point of each driving wheel i with the x-axis parallel to the rotation axis of each wheel, whose origin positions rri \u2208 R 3 are function of the radius R of the circle in which each wheel is placed equidistantly from each other, and the vertical position h of the local frame w" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000975_fit.2013.43-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000975_fit.2013.43-Figure1-1.png", + "caption": "Figure 1: The Swerve Drive", + "texts": [], + "surrounding_texts": [ + "Keywords\u2014Velocity control ,Omni-directionalrobot, Swerve Drive, PI controller, Robotics, Proportional, Integral.\nI. INTRODUCTION DC motors are widely used in the industry because of the reliability, robustness and the effective control algorithms. Velocity control of a DC motor is one of the most important aspects in industrial control systems. Almost all industrial automation equipment, Electric cars and industrial/military robots are equipped with DC Motors as actuators. Speed of the motor depends on many factors including voltage, load and temperature. Almost all of the motor applications need the velocity to remain constant at a certain level even if above mentioned parameters change. These algorithms can control the angular velocity of the motors according to the needs. Most of these control methods are implemented on automated machinery ranging from simple Packaging systems to Complex Military Robots. In order to navigate a four wheeled mobile robot in straight direction, all the four wheels should have exactly same speed. If any one of the wheels has different speed it would make the robot to deviate from the straight direction. It is impossible to maintain exactly same speed on each wheel, without using any feedback mechanism. The feedback can be taken by a transducer whose output is a function of motor speed. The two main issues encountered without using this technique are: (1): Motor speed is directly proportional to the applied Voltage across it. The speed varies with the variation in the applied voltage level [1]. During the operation of robot, it consumes power from the battery which causes the Voltage level of the battery to drop (depending upon how much power is consumed). The voltage applied across the motor also drops and thus the speed of the motor decreases. On the other hand, fully charged battery will cause the speed of motor to increase.\n(2): Secondly, the speed of motor is also affected by the\namount of weight placed upon the robot and by the distribution of the weight placed upon the robot [2]. Speed of the robot would be greater when unloaded than when loaded. When load is increased on the motor more torque is required to maintain the same speed. On the other hand if load is decreased the speed of the motor will increase under the same conditions [3]. To overcome these issues we developed and implemented a feedback based motor control algorithm which makes the speed of the motors of a mobile robot independent of the above mentioned problems. But there are still some mechanical and electrical constraints that affect the algorithm. Those constraints are addressed in Section V.\n978-1-4799-2293-2/13 $31.00 \u00a9 2013 IEEE\nDOI 10.1109/FIT.2013.43\n195", + "The sensor used to take feedback from the m Incremental Quadrature encoder. The output of t is square wave whose frequency depends upon th the motor. The quadrature encoder used ha (pulses-per-revolution) i.e. each motor increments/decrements the counts value by 25 output of quadrature encoder (coupled to the s motor) is decoded and then its data is used to id much the motor is moved[6].\nTo drive the Robot efficiently and swiftly Bru Motors (BLDC) were used in the robot. As suggests, BLDC motor lack commutation brush of brushes commutation is done electronically motor has permanent magnets as Rotor which windings as Stator, eliminating the problems of current via mechanical contact to the moving Due to the absence of brushes, BLDC motor pro Electromagnetic Interference and has high powe ratio. The lack of physical contact for co purpose also makes BLDC motor more efficien of power dissipation and increases its life compared to a DC Brushed motor[7].These BLDC motor make it a perfect actuator to b Industrial automation equipment and in Robotics\nMotor drives are electronic power amplifiers th to drive the motors.These drive circuits requir signal to drive the motor. The speed of the mot upon the width of this control signal, this controlling speed is commonly known as Pu Modulation. This electronic speed controller i from a DC power source and it is a 3-phase In generates a 3-phase quasi square wave which is Brushless DC Motor, for its operation.\nIII. ALGORITHM IMPLEMENTATIO The algorithm developed and implemented is b Proportional Integral (PI) Controller. PI cont mostly used in industrial applications. Th excluding the Derivative term \u2018D\u2019 is usually us fast response of the system is not required a when disturbances and noise affect the overall especially the microcontroller operation. The time selected for the PI-controller is 10milisec Sampling time is selected while keeping in min resolution at worst case scenario and smooth between throttle levels. The Proportional and Int are selected by taking account of the Ziegler-N tuning criteria [9]. A value of \u2018desired counts\u2019 to the PI-controller. After each sampling interv calculated between the desired counts and the ac incurred. The actual counts are basically the c result due to the movement of the motor in t 10miliseconds (sampling interval). Figure.2 complete system model where as Figure.3 is the of the Velocity Controller that is implemen Controller block in Figure.2.\notor is an he encoder e speed of s 250PPR revolution 0 [5]. The haft of the entify how\nshless DC the name es, instead . A BLDC rotate and providing armature. duces less r to weight mmutation t in terms period, as features of e used in .\nat are used e a control or depends method of lse Width s powered verter that fed to the\nN ased upon rollers are e control ed when a nd mainly system [8], sampling onds. The d the error transition egral gains icholas PI is inputted al, Error is tual counts ounts that he time of shows the Flow chart ted in the\nIf Error resulted is Positive then the increased by a factor. In case of Neg the motor is decreased by a factor. T the amount of error occurred; the lar the factor and vice versa. After calcu is fed into the Electronic Speed C drives the BLDC motor.\nspeed of the motor is ative Error, the speed of he factor depends upon ger the error, the greater lations the control signal ontroller, which further\nodel\ner Flow Chart" + ] + }, + { + "image_filename": "designv11_34_0002954_9781119195740-Figure9.16-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002954_9781119195740-Figure9.16-1.png", + "caption": "Figure 9.16 An elbow manipulator in its side and top views.", + "texts": [ + " (b) Second Kind of Motion Singularity Equations (9.168) and (9.169) imply that the second kind of motion singularity occurs if s\ud835\udf035 = 0, i.e. if \ud835\udf035 = 0 or \ud835\udf035 =\ud835\udf0b. However, while this singularity can occur easily with \ud835\udf035 = 0, it can hardly occur with \ud835\udf035 = \ud835\udf0b due to the physical shapes of the relevant links and joints. In other words, this singularity is the same as the third kind of motion singularity of the Puma manipulator, which is explained in Section 9.1.5 and illustrated in Figure 9.11. An elbow manipulator is shown in Figure 9.16. The first example of manipulators with this generic name is the Cincinnati Milacron T3 Robotic Arm developed by Richard Hohn for the Cincinnati Milacron Corporation. Actually, the class of elbow manipulators is more inclusive than the group of manipulators like the one depicted in Figure 9.16. It covers all the articulated manipulators with anthropomorphic arms. An articulated manipulator comprises revolute joints only and an anthropomorphic arm possesses an elbow joint. In this general sense, a Puma manipulator is also an elbow manipulator. However, since the Cincinnati Milacron T3 is the first anthropomorphic robot with an elbow joint, it is considered to be the namer for the class of elbow manipulators. An elbow manipulator comprises six revolute joints. So, it is designated symbolically as 6R or R6. Its joint axes are also shown in Figure 9.16 together with the relevant unit vectors. The significant points of the manipulator are named as follows: O: Center point (origin of the base frame) E: Elbow point, Q: Pre-wrist point, R: Wrist point, P: Tip point The shoulder point of this manipulator is coincident with its center point, i.e. S = O. Differently from the Puma and Stanford manipulators, an elbow manipulator has a noticeable feature that its wrist is not spherical, i.e. the axes of the last three joints do not intersect at a single point. (a) Joint Variables \ud835\udf031, \ud835\udf032, \ud835\udf033, \ud835\udf034, \ud835\udf035, \ud835\udf036 They are the rotation angles about the joint axes. They are shown in Figure 9.16. (b) Twist Angles \ud835\udefd1 = 0, \ud835\udefd2 = \ud835\udf0b\u22152, \ud835\udefd3 = 0, \ud835\udefd4 = 0, \ud835\udefd5 = \u2212\ud835\udf0b\u22152, \ud835\udefd6 = \ud835\udf0b\u22152 (c) Offsets d1 = 0, d2 = 0, d3 = 0, d4 = 0, d5 = 0, d6 = RP (tip point offset) (d) Effective Link Lengths b0 = 0, b1 = 0, b2 = OE, b3 = EQ, b4 = QR, b5 = 0 (e) Link Frame Origins O0 = O,O1 = O,O2 = O,O3 = E,O4 = Q,O5 = R,O6 = P 260 Kinematics of General Spatial Mechanical Systems (a) Link-to-Link Orientation Matrices C\u0302(0,1) = eu\u03031\ud835\udefd1 eu\u03033\ud835\udf031 = eu\u03033\ud835\udf031 (9.172) C\u0302(1,2) = eu\u03031\ud835\udefd2 eu\u03033\ud835\udf032 = eu\u03031\ud835\udf0b\u22152eu\u03033\ud835\udf032 (9.173) C\u0302(2,3) = eu\u03031\ud835\udefd3 eu\u03033\ud835\udf033 = eu\u03033\ud835\udf033 (9", + "187) can be manipulated as shown below. r = b2C\u0302(0,2)u(2\u22152) 1 + b3C\u0302(0,3)u(3\u22153) 1 + b4C\u0302(0,4)u(4\u22154) 1 \u21d2 r = b2eu\u03033\ud835\udf031 e\u2212u\u03032\ud835\udf032 eu\u03031\ud835\udf0b\u22152u1 + b3eu\u03033\ud835\udf031 e\u2212u\u03032\ud835\udf0323 eu\u03031\ud835\udf0b\u22152u1 + b4eu\u03033\ud835\udf031 e\u2212u\u03032\ud835\udf03234 eu\u03031\ud835\udf0b\u22152u1 \u21d2 r = b2eu\u03033\ud835\udf031 e\u2212u\u03032\ud835\udf032 u1 + b3eu\u03033\ud835\udf031 e\u2212u\u03032\ud835\udf0323 u1 + b4eu\u03033\ud835\udf031 e\u2212u\u03032\ud835\udf03234 u1 \u21d2 r = eu\u03033\ud835\udf031(b2e\u2212u\u03032\ud835\udf032 u1 + b3e\u2212u\u03032\ud835\udf0323 u1 + b4e\u2212u\u03032\ud835\udf03234 u1) (9.188) Equation (9.188) can also be written as r = eu\u03033\ud835\udf031 r\u2032 (9.189) In Eq. (9.189), r\u2032 = r(1). That is, r\u2032 = u1(b2c\ud835\udf032 + b3c\ud835\udf0323 + b4c\ud835\udf03234) + u3(b2s\ud835\udf032 + b3s\ud835\udf0323 + b4s\ud835\udf03234) (9.190) Note that, referring to Figure 9.16, r\u2032 = u1r\u20321 + u2r\u20322 + u3r\u20323 can also be written down directly by inspection. In other words, Figure 9.16 readily implies the following coordinates of the wrist point R in the frame 1(O). r\u20321 = b2c\ud835\udf032 + b3c\ud835\udf0323 + b4c\ud835\udf03234 r\u20322 = 0 r\u20323 = b2s\ud835\udf032 + b3s\ud835\udf0323 + b4s\ud835\udf03234 \u23ab \u23aa\u23ac\u23aa\u23ad (9.191) On the other hand, the expansion of the expression in Eq. (9.188) leads to the coordinates of the wrist point R in the base frame 0(O) as follows: r(0) = r = u1r1 + u2r2 + u3r3 (9.192) r = (eu\u03033\ud835\udf031 u1)r\u20321 + (eu\u03033\ud835\udf031 u3)r\u20323 \u21d2 r = (u1c\ud835\udf031 + u2s\ud835\udf031)r\u20321 + (u3)r\u20323 \u21d2 r = u1(r\u20321c\ud835\udf031) + u2(r\u20321s\ud835\udf031) + u3(r\u20323) (9.193) Equations (9.193) and (9.191) indicate that r1 = (b2c\ud835\udf032 + b3c\ud835\udf0323 + b4c\ud835\udf03234)c\ud835\udf031 r2 = (b2c\ud835\udf032 + b3c\ud835\udf0323 + b4c\ud835\udf03234)s\ud835\udf031 r3 = b2s\ud835\udf032 + b3s\ud835\udf0323 + b4s\ud835\udf03234 \u23ab \u23aa\u23ac\u23aa\u23ad (9", + " On the other hand, if the task to be executed necessitates keeping the manipulator in such a pose, then the freedom in \ud835\udf032 may be used to orient the folded arm links of the manipulator conveniently depending on the environmental conditions. For example, the coincident upper and front arms may be allowed to get aligned with the gravity direction so that the task can be executed with minimal potential energy. 268 Kinematics of General Spatial Mechanical Systems (c) Third Kind of Position Singularity Equations (9.212) and (9.213) imply that the third kind of position singularity occurs if s\ud835\udf035 = 0, i.e. if \ud835\udf035 = 0 or \ud835\udf035 = \ud835\udf0b. In other words, as also implied by Figure 9.16, this singularity occurs if the orientation of the end-effector is specified so that the approach vector is perpendicular to the front arm. This singularity is illustrated in Figure 9.22 with both \ud835\udf035 = 0 and \ud835\udf035 = \ud835\udf0b. If this singularity occurs, \ud835\udf03234 and \ud835\udf036 become indefinite so that they cannot be found separately. This is because they turn out to be rotation angles about parallel axes, which are the axis of the sixth joint and the already parallel axes of the fourth, third, and second joints. Therefore, the angles \ud835\udf03234 and \ud835\udf036 cannot be distinguished from each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001482_1.4029187-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001482_1.4029187-Figure2-1.png", + "caption": "Fig. 2 Equivalent four-bar linkage Fig. 3 A dead center position of Stephenson type II linkage", + "texts": [ + " Any of the equivalent four-bar linkages of the single-DOF planar linkage is at the dead center positions, the whole linkage must be at the dead center positions. These conclusions can be explained as following. Let link 1 be the reference link, link 2 be the input link, link p and link q be any other two links of the complex linkage. Thus, the corresponding 044501-2 / Vol. 7, NOVEMBER 2015 Transactions of the ASME Downloaded From: http://mechanismsrobotics.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use instant centers I12, I2p, Ipq, and Iq1 form an equivalent four-bar linkage, as shown in Fig. 2. The velocity of link p and link q relative to link 2 can be expressed with the instant centers as wp w2 \u00bc I12I2p I1pI2p (1) wq w2 \u00bc I12I2q I1qI2q (2) where wi (i\u00bc 2, p, q) is the angle velocity of link i. For the complex linkage staying at the dead center position, Eqs. (1) and (2) should be infinite. Hence, the denominators of the right side of these two equations can be zero. That is, I1pI2p \u00bc 0 and I1qI2q \u00bc 0. This is to say the instant center I1j and I2j (j\u00bc p, q) coincide. According to Aronhold\u2013Kennedy\u2019s theorem, the instant center Ipq should be on the line I1pI1q and on the line I2pI2q" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003114_042079-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003114_042079-Figure1-1.png", + "caption": "Figure 1. Simplified simulation model of the tractor MTZ-82.1.", + "texts": [ + " ICMSIT 2020 Journal of Physics: Conference Series 1515 (2020) 042079 IOP Publishing doi:10.1088/1742-6596/1515/4/042079 As an object of modeling and research, a machine-tractor unit based on the MTZ-82.1 tractor was selected. It was equipped with mounted modular implements [16]. Modeling was performed in CAD SolidWorks and CAE SolidWorks Motion by the method of multi-body dynamics (MBD). At the first stage of the study, a model of a tractor equipped with front and rear three-point linkage was created (figure 1). All the basic structural elements and their geometric and mass-inertial parameters were preserved. All stationary elements of the tractor were excluded from the model (shown in the figure in wireframe). To replace them, we used an equal mass ball of custom high density material. By adjusting the position of the ball, the center of mass of the simplified model is combined with the real operational center of mass of the tractor. This simplification allows to perform simulation with minimal computer load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002198_978-3-319-14194-7_4-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002198_978-3-319-14194-7_4-Figure5-1.png", + "caption": "Fig. 5 Two-link planar manipulator", + "texts": [ + " Then (41) can be written as _r\u00feM 1\u00f0Kv \u00feVm \u00fe 1\u00der \u00bc M 1\u00f0F\u00fe sd\u00de \u00f043\u00de The dynamical system (43) is stable as long as the friction term and disturbances are bounded. This means filtered error r(t) remains bounded and thus the tracking error e\u00f0t\u00de and its time-derivative _e\u00f0t\u00de are also bounded. The proposed trajectory tracking control scheme with the adaptive P-T2FCMAC friction and disturbance compensator is shown in Fig. 4. The performance of the proposed adaptive P-T2FCMAC neural compensator is evaluated when applied on a two-link RR planar manipulator shown in Fig. 5. It has been modeled as two rigid links of length 1 m each with masses m1 \u00bc 1 kg and m2 \u00bc 1 kg concentrated at the distal ends of the links. The following nonlinear viscous and dynamic friction terms of F\u00f0 _q\u00de and unknown disturbances \u03c4d have been included in the manipulator dynamics. F\u00f0 _q\u00de \u00bc Fv _q\u00feFd sign\u00f0 _q\u00de sd \u00bc Td sin\u00f015t\u00de \u00f044\u00de where Fv = 1; Fd = 0.5; Td = 0.8. The desired trajectory to be followed is given by qd1\u00f0t\u00de \u00bc sin 2pt=T\u00f0 \u00de\u00fe 1; qd2 \u00bc cos 2pt=T\u00f0 \u00de\u00fe 1 \u00f045\u00de where T = 2 s. The matrix Kv is taken as Kv = [100 0; 0 20]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002483_itec-india.2015.7386862-Figure8-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002483_itec-india.2015.7386862-Figure8-1.png", + "caption": "Figure 8. Electrical \u2013 Turbo Compounding Engine System", + "texts": [], + "surrounding_texts": [ + "Gas turbines have been previously used and researched by America (M1 Abrahms) and Russia (T \u2013 80) unsuccessfully, but as it was used as a direct energy source its shortcomings were highlighted more than its advantages." + ] + }, + { + "image_filename": "designv11_34_0003626_ilt-04-2020-0144-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003626_ilt-04-2020-0144-Figure1-1.png", + "caption": "Figure 1 Schematic position of defects", + "texts": [ + " Literature review reveals that majority of the researchers focused on bearing-fault identification, dynamic behavior of rolling bearing due to localized defects and the theoretical research of rotor system support by ball bearings with SFD or ISFD. However, literature survey also indicates a dearth of the experimental research papers pertaining to vibration studies of rotor system support by defective ball bearings with ISFD. ISFD can be adopted as an effective way to improve security and stability. Therefore, the purpose of this paper is to experimentally investigate the effects of ISFD on the dynamic characteristics of the faults rotor system. Defects on the surface of races and ball have been shown schematically in Figure 1. According to Tadina and Bolte\u017ear\u2019s (2011) work, defect on races was modeled as an ellipsoidal depression, and defect on ball was modeled as a flattened sphere. It can reflect the actual path of ball in a faulty bearing. The ball has an additional deflection whenever the defect is in contact with its mating surface; this would lead to the change of Hertz contact force. During contacting, balls periodical impact with races due to defects generate periodical impulses which are affected by defect characteristics, duration and relative speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001131_icrom.2013.6510116-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001131_icrom.2013.6510116-Figure2-1.png", + "caption": "Figure 2 -Parameters definition of the assumed two WMM", + "texts": [], + "surrounding_texts": [ + "The dynamics and kinematics model of a two wheel mobile manipulator robot with a reaction wheel is described in this section. For the considered robot which is shown in Figures (1-2), its parameters specifications are given in Table (1). The base of the mobile manipulator is assumed as a passive joint. Therefore, the manipulator part consists of three links and actuators in 3D workspace. The reaction wheel is placed on the first link which has one more degree of freedom rather than this one. The dynamics equations of motion of a mobile manipulator are described as: '( = M(q)q + H(q,q) + G(q) (1) where q=[ql'q2'\" .. ,qtER7 is the vector of generalized coordinates and =[\" 2,0, 4' 5' 6' lER7 is the input generalized torque and M(q) ER7x7, H(q, q) ER7xI, G( q) ER7X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively. The considered two WMM robot with a reaction wheel required complex dynamics modelling as a result of under actuated system. The passive joint causes that the balancing challenge of this robot in the XOY plane is more important than other its positions. Thus, Double Inverted Pendulum Model (DIPM) is utilized to simplify the dynamic analysis for the balancing control, [18]. The effectiveness of this simplified model is demonstrated for the dynamic locomotion of the highly nonlinear and complex system, [19]. Tablel - Parameters of the two WMM with a reaction wheel o XYZ OI_XIYIZI Q,&q2 Q3 Q;(i=4,5,6) q, L 1,(i= I ,2,3) 14 Word coordinate frame Mobile manipulator coordinate frame Rotation angles of wheels Inclination angle of the passive joint Joint angles of links Angular position of the reaction wheel Radius of wheels Distance between wheels Length of links Distance of the reaction wheel centre to first joint B. Dynamics Modelling of Double Inverted Pendulum It is possible to model two WMM as a virtual double inverted pendulum model. In this model, the components of second and third of manipulator in XOY plane are considered as the virtual second link of double inverted pendulum model. This simple model is used as the stabilizing control part. This model is shown in Figure (3) and parameters of these models is stated in Table (2). Table2 - Parameters of DIPM ilw Rotation angle of wheels ill Joint angle of the first link ile Joint angle of the second link il2 Angular position of the reaction wheel 11 Length of the first link Ie Length of the second I ink 12 Distance of reaction wheel centre to first joint xme X position of the me in 01 coordinate frame Yme Y position of the me in 01 coordinate frame Xl X component of the first link Length Y1 Y component of the first link Length me Equivalent mass on the second link m1 Mass of the first link m2 Mass of the reaction wheel mw Mass of wheels c,(i=2,e) Coefficient of friction a1 Length of the CoG on the first link i1 Moment inertia of link I around the ZI axis i2 Moment inertia of the reaction wheel around the zraxis Calculation of positions of the virtual link xmeand Yme' and its length Ie' joint angleqe and the mass meare as following: (2) (3) (4) (5) The dynamics equations of motion of this model are obtained by Euler-Lagrange equation as: (6) whereq = [qw, q1' q2' qeF ER4 is the vector of generalized coordinates and r = [fw, f1' f2' feF ER4 is the input generalized torque for the virtual double inverted pendulum modeI.M(q) ER4X4,A(q, q) ER4XI,G(q) ER4X1 are the inertia matrix, centrifugal and Coriolis terms, and the gravity matrices, respectively which are calculated as: M = [\ufffd\ufffd\ufffd M31 M41 where: Mll = (mw + m1 + mz + me)rZM12 = M21 = (a1m1 + me (11 + Ie COS(qe)) + 11mz)rcOS(q1) + mwrZM14 = M41 = melerCOS(q1 + qe) Mzz = i1 + iz + a1 Z m1 + 11 z mz + 11 z me + Ie z me cosz qe + 211 leme cos qe + mWrZM23 = M32 = M33 = izMz4 = Ie z me cosz qe + 11 leme cos qe M13 = M31 = M34 , Z = M43 = OM44 = Ie meH1 = -fiI(mer(ft1 sin(i'iI) (\\1 + Ie COS(qe)) + leqe sin(cl!+qe)) + 11mZft1rsin(Q1) + a1 m1 ib r Sin(Q1)) - merqe(leqe Sin(q1 + qe) + leq1 COS(q1) sin(qe))Hz = -2m.leq1 (feqe COS(qe) sin(qe) + 11 qe sin(qe)) + meleqe (qwr COS(q1) COS(qe) - qe sin(qe) (\\1 + 21e cos(qe))) H3 = cZqZH4 = -me Ie rih qw COS(qe) sin(qe) + m.leq1 z sin(qe) (11 + Ie COS(qe)) + CeqeG1 = G3 = OGz = -a1 m1 Sin(q1) - 11 sin q1 (mz + me) - Ie me (cos q1 sin qe + sin q1 cos qe)G4 = -Ieme sin(q1 + qe) C. Dynamic Verification/or DIPM ADAMS is the most widely used multi-body kinematics and dynamics analysis software in the word. Also, Adams helps engineers to study the kinematics and dynamics of moving parts, how loads and forces are distributed throughout mechanical systems, and to improve and optimize the performance of their products. The extracted dynamics equations of motion of the virtual double inverted pendulum with a reaction wheel are verified by the ADAMS model which the 3D sketch of DIPM is shown in Figure (4).We insert some similar torque as input to these two models and compare the reaction of model joints with together. This torques is shown in Figure (5) which insert to the reaction wheel. Moreover, the obtained result of the model response is shown in Figure (6).These curves show that the dynamics equations of motion of the virtual double inverted pendulum are verified. So, we can use these equations in the control algorithm to realize the dynamic stability. In addition, the parameters specifications of this system in this verification routine are expressed in Table (3). 0.4 0.3 0.2 E 0.1 ;;. ! -0.1 -0.2 -0.3 I I I I I I I I I ___ 1 ____ 1___ -.l ___ .l ___ L ___ 1 __ _ 1___ -.J ___ \ufffd __ _ I I I I -- \ufffd --- 1 --- T --- \ufffd -- -0.4 - - -1- - - -1 - - - -+ - - - + - - - t- - - - 1- - - -1- - - ---j I I I I I I I I -0.5 OL--- :0c'c .5=- -L.- 1\ufffd.5=- -\ufffd2 --2cc'.C:-5-\ufffd3---=-3.L5 --':-\ufffd4\"'.5-\ufffd 5 Time (8) Since, the two WMM contains a passive joint, it requires an active controller to stabilize the passive joint and to control its stability. Related to ZMP criterion,[20], this robot have just one stable natural position that the moment of the gravity has no effect on the body, in other words, gravity force direction of COG crosses the mobile wheel axis. If we control the position of the COG to this point, we can able to satisfy the robot stability. In this section, a PID controller is proposed to control the motion of the reaction wheel to achieve dynamic stability. The supervising controller identifies the position of the COG and tunes the set point of PID controller. At the same time, the PID controller moves the reaction wheel to reach the stable natural position. Also, the control block diagram of this controller is shown in Figure (7). The supervising controller finds the COG position without any information about robot links and the related body. So, as shown in Figure (7), the input parameters of this block are the reaction wheel angular velocity and its acceleration. Therefore, it can just find the position of the COG to the right or the left from the normal position of the first link (qi=O) with this information. For an example, in Figure (2), if the reaction wheel angular velocity and its acceleration are positive, it is clear that the position of COG are in the left side from the normal position of the first link (qi=O).SO, the supervising controller changes the PID set point to the negative value. Thus, this can use many type of controller for this block such as PI, PID, Fuzzy or GA controller. In this paper we used the PI controller for this block. IV. SIMULATION RESULTS AND DISCUSSION The validation of the proposed control strategy is demonstrated by simulation results. This simulation runs at MA TLAB Simulink Toolbox for the assumed robotic system which the parameters specifications are expressed in Table(3). In addition, the controller gains are expressed in Table (4). Table4--Controller gains amount Kp Coefficient of P action 3.916 KI Coefficient of I action 1 KD Coefficient of 0 action 0.0624 Ksd Coefficient of q3 parameter -6e-3 Ksp Coefficient of q3 parameter 0.0209 This type of robot is dynamically stable and this stability is attained using a reaction wheel. Here, we consider the three bench marks and run the simulations. In the first and second one, the robot start from an initial position and make itself stable and, in third one, the robot run in the different initial position and pass the variable acceleration. The initial conditions of these case studies are respectively expressed in Table (5). Moreover, the accelerate variation curve of the third case is shown on Figure (8). Table 5 - Initial conditions for simulations Casel Casell CaselII q qp i'i2 Joint angle of the first link (deg) -30 -30 30 Joint angle of the second link (deg) -60 -120 90 Angular position of the reaction wheel (deg) 0 0 0 1.5 - - - - - I - -----1- - - - - - 1 - - --- 1 ----- - - - - - ---1 - -----1- - - - - - r- - ---- _____ --J ______ 1 ______ L ____ _ ........ 0.5 1 .\ufffd \ufffd ] -0.5 I I I ----- T -----1-----l------------ \ufffd -----r----- \ufffd----- \ufffd------------ \ufffd.5 -----\ufffd----- 4----- \ufffd------------ -20\ufffd------\ufffd,0\ufffd----\ufffd2\ufffd0 ----\ufffd3\ufffd 0------\ufffd4 0\ufffd----\ufffd 5\ufffd 0------\ufffd6 0 Time(s) Figure 8 - Acceleration variation during the operation As shown from simulation results in Figures (9-17), the time history of the robot motion is presented in Figures (9- 11) for the cases to move and reach to the stable position. Moreover, the position of the <11 must be varied until the CoG reaches to the appropriate position Figures (12-14). In this position, the robot is stable and there is no moment appears to change this style. Also, the PID modifies the set point value to satisfy the stabilization. Figures (15-17) are shown the position of the reaction wheel, during the robot reaches to stable position. These results show that the reaction wheel movement makes the internal moment that it can control the passive joint. In this control strategy, the balancing is not achieved by the robot movement and only the reaction wheel is used to stabilize. This method can also improve the dexterity of the robot motion and increases the robot controllability." + ] + }, + { + "image_filename": "designv11_34_0001593_1.4032132-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001593_1.4032132-Figure1-1.png", + "caption": "Fig. 1 Electropneumatic system", + "texts": [ + " The adaptive version is similar to the first observer with the difference that the observer design parameter is no longer constant but is governed by a Riccatti differential equation, and the involved adaptation process is mainly driven by the power of the output observation error norm computed on a moving horizon window. This paper is organized as follows: In Sec. 2, we present the global electropneumatic system model. Section 3 is devoted to describe the class of systems under investigation and to synthesize the classic and dynamic high-gain observers. In Sec. 4, we apply those observers to electropneumatic system and we present simulation results to compare between the two observers. In Sec. 5, some concluding remarks are provided. The considered system (Fig. 1) is a linear double acting actuator composed by two chambers denoted P (positive) and N (negative). Two three-way servodistributors supply the air mass flow rates entering to the two chambers. The actuator rod is connected to one side of the carriage and drives an inertial load on guiding rails. The total moving mass is 17 kg. Table 1 shows the specifications of the pneumatic actuator. 2.1 Mechanical Model. According to Newton\u2019s second law applied to the moving solid and under the assumptions of a link rod/rigid carriage and a rigid body, the dynamics equation of movement is dy dt \u00bc v dv dt \u00bc 1 M SPpP SNpN bvv Fext\u00bd 8>< >: (1) where y and v denote the position and the velocity of the rod, respectively; pP and pN, respectively, stand for the pressure in chambers P and N; bv is the viscosity coefficient; and Fext is an external constant force due to the atmospheric pressure Patm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000924_chicc.2014.6895471-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000924_chicc.2014.6895471-Figure1-1.png", + "caption": "Fig. 1: Quad-rotor helicopter scheme", + "texts": [ + " In section 3, first, the internal model principle is introduced to deal with the tracking problems, then the robust adaptive H\u221e fault-tolerant tracking controller is designed based on the fault model. In section 4, a numerical of 3D hover system model and its simulation results are given. Assuming the quad-rotor structure is a rigid body and symmetrical, the center of the mass and the body fixed frame origin are coincide, and the thrust and drag torque is proportional to the square of the the propeller speed same as the reference [17, 18]. Let B = {Bx, By, Bz} denote the body fixed frame and E = {Ex, Ex, Ez} denote the earth fixed frame as figure 1. The vector \u03b7 = [\u03c8,\u03c6, \u03b8] denote the attitude angles of the quad-rotor relative to the body fixed frame. The attitude angles (Euler angles) {\u03c8,\u03c6, \u03b8} are respectively called roll angle (\u2212\u03c0 2 < \u03b8 < \u03c0 2 ), and yaw angle (\u2212\u03c0 < \u03c8 < \u03c0), pitch angle (\u2212\u03c0 2 < \u03c6 < \u03c0 2 ). The body X, Y, Z axis are defined as the roll axis, pitch axis and yaw axis, respectively. The roll angle \u03b8, the pitch angle \u03c6 and the yaw angle \u03c8 are defined as Fig.1. Assuming the pitch angle or the roll angle are zero when we use the Euler equation to get the roll angle and pitch angle motion equation. Then the roll angle and pitch angle equations are as follows: Ixx\u03b8\u0308 = \u03c4x Iyy\u03c6\u0308 = \u03c4y (1) The yaw angle is defined as positive for a counterclockwise rotation (when looking down on the system from above) and drive by motor 1,3 and motor 2,4. Then Yaw angle motion equation is that: Izz\u03c8\u0308 = \u03c4z (2) The torque that control quad-rotor helicopter attitude an- gle respectively is described as follows: \u03c4x = bl(\u03a92 2 \u2212\u03a94 2) \u03c4y = bl(\u03a91 2 \u2212\u03a93 2) \u03c4z = d(\u03a94 2 + \u03a92 2 \u2212\u03a93 2 \u2212\u03a91 2) (3) The physical meaning of above symbols are described as Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001769_robio.2015.7418980-Figure6-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001769_robio.2015.7418980-Figure6-1.png", + "caption": "Fig. 6 Parameters of D-H model", + "texts": [ + " Firstly, the kinematic model of industrial robot is established by using the D-H modeling method[8], then analyze the forward kinematics and inverse kinematics. By using a laser tracker, the calibration model of robot was built with the Gauss-Newton method. A. Kinematic Model of Robot Before the kinematics analysis, robot\u2019s kinematic model should be established. Industrial robots are generally link mechanism, which is composed of rigid links connected by joints. The main parameters of D-H model are showed in Fig. 6 including length of link , twist angle , setover , joint angle . According to the theory of rigid body translation and rotation, the transformation matrix between adjacent coordinates can be obtained. , 0,0, ,0,0 x, 1 0 0 0 0 0 0 1 0 0 0 0 0 0 1 0 0 1 0 0 0 1 0 0 0 1 1 0 0 1 0 0 0 0 1 0 0 0 0 \u00a0 0 0 1 0 0 0 0 0 0 1 0 0 0 1 i i i i i i i i i i i i i i i i i i i i i i i i A Rot z Trans d Trans a Rot c s s c d d a c s s c c s c s s a c s c 0 0 0 0 1 i i i i i i i i c c s s s c d (10) Where \u00a0 , c is cos s is sin " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000409_aero.2015.7118995-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000409_aero.2015.7118995-Figure3-1.png", + "caption": "Figure 3. Cylindrical roller bearing and equivalent model of two cylinders. Surface compressive stress distribution is represented by ideal line contact [7].", + "texts": [ + " The damage initiation process was simulated using the elastic damage model based on the work published by Slack and Sadeghi [3] and by Warhadpande et al. [4]. A discussion of the obtained results concludes the article. While rolling element bearings are loaded, a small contact area between the rolling element and the raceway is generated. The Hertzian contact representation was selected assuming that during most of the time, the rolling element bearings are loaded below the yield stress limit. This section presents a simulation of static and rolling contact cases. An ideal line contact, representing the cylindrical roller bearing (Fig. 3), is used in both cases. The static contact is presented for the elastic case only, while the rolling contact is presented for two load regimes, elastic and elastic-plastic. The surface compressive stress distribution, p, within the line contact area is given by p = pmax ( 1\u2212 x2 b2 )0.5 (1) were pmax, b and x are maximum pressure at the middle of the pressure profile, half width of the contact and local coordinate, respectively. An area, representing contact between the roller and the raceway, (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003802_s12541-020-00438-1-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003802_s12541-020-00438-1-Figure10-1.png", + "caption": "Fig. 10 Limitation of local ZMP in the foot with a safety margin for preventing rolling over", + "texts": [ + "\u00a0(11) is an inequality condition that prevents the local ZMP of each sole from crossing the edge of each sole to prevent the sole from rolling over the ground. lL,f , lR,f are the lengths between the front edge and ankle of the left and right foot, lL,b, lR,b are the lengths between the rear edge and ankle of the left and right foot, lL,L, lR,L are the lengths between the left edge and ankle of the left and right foot, and lL,R, lR,R are the lengths between the right edge and ankle of the left and right foot, respectively. Each length was multiplied by 0.8 to give a 20% margin (see Fig.\u00a010). where , Q = diag(1, 1, 1), S = diag(0.01, 0.01, 1, 1, 1, 1) , C = diag(1, 1, 1, 1, 1, 1), flower = [ F QP z,L,l F QP z,R,l M QP Pitch,L,l M QP Pitch,R,l M QP Roll,L,l M QP Roll,R,l ]T , (9)min f (Af \u2212 b)TQ(Af \u2212 b) + f TSf (10)flower < Cf < fupper (11)0.8 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 lL,bF QP z,L lR,bF QP z,R lL,LF QP z,L lR,LF QP z,R \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 < \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 M QP Pitch,L M QP Pitch,R M QP Roll,L M QP Roll,R \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 < 0.8 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a3 lL,f F QP z,L lR,f F QP z,R lL,RF QP z,L lR,RF QP z,R \u23a4\u23a5\u23a5\u23a5\u23a5\u23a6 f = [ F QP z,L F QP z,R M QP Pitch,L M QP Pitch,R M QP Roll,L M QP Roll,R ]T ADS = \u23a1\u23a2\u23a2\u23a3 1 1 0 0 0 0 xL \u2212 p ref x xR \u2212 p ref x 1 1 0 0 yL \u2212 p ref y yR \u2212 p ref y 0 0 1 1 \u23a4\u23a5\u23a5\u23a6 ALSS = \u23a1\u23a2\u23a2\u23a3 1 0 0 0 0 0 xL \u2212 p ref x 0 1 0 0 0 yL \u2212 p ref y 0 0 0 1 0 \u23a4 \u23a5\u23a5\u23a6 ARSS = \u23a1\u23a2\u23a2\u23a3 0 1 0 0 0 0 0 xR \u2212 p ref x 0 1 0 0 0 yR \u2212 p ref y 0 0 0 1 \u23a4\u23a5\u23a5\u23a6 b = [ Fref z 00 ]T , 1 3 In this study, a force-position hybrid control is used to control the balance of the robot as shown in the block diagram in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003416_aim43001.2020.9158920-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003416_aim43001.2020.9158920-Figure1-1.png", + "caption": "Fig. 1: Block diagram and real-world setup of the inverted pendulum with dual-axis reaction wheels.", + "texts": [ + " This work was partially supported by the Nazarbayev University Faculty Development Competitive Research Grant no. 090118FD5339, \u201cHardware and Software Based Methods for Safe Human-Robot Interaction with Variable Impedance Robots\u201d. physical interaction is inherently not possible. On the other hand, thanks to their high actuation bandwidth, the reaction wheels will be able to correct small amplitude deviations from the desired trajectories due to model uncertainty or external disturbances. In order to test this hypothesis, we created a compliant dual-axis pendulum setup actuated by reaction wheels (see Fig. 1). Controlling such a system requires planning the motion of the reaction wheels within a certain time horizon in the future (typically, a few seconds) so as to bring the system to a final rest configuration. The natural solution for such a task would be to use optimal control, ideally solving an optimal control problem (OCP) over an infinite time horizon [4]. Explicit solutions of such a problem can be obtained using the linear quadratic regulator (LQR) approach, which however is possible only for linear systems, and when no inequality constraints on input or state variables are present [5]", + " Section III introduces the general OCP formulation, together with its approximations via DNN and NMPC. Section IV introduces the experimental setup, including the hardware components and the procedures employed to obtain the system parameters; furthermore, it provides a description of the implementation and tuning of the DNN and NMPC schemes. Finally, Section V presents and discusses simulation and experimental results, and conclusions are drawn in Section VI. The block diagram of the inverted pendulum with dualaxis reaction wheels is shown in Fig. 1a. The rotational motion of the reaction wheels about their center of mass changes the angular momentum, which rotates the pendulum in the opposite direction [13]. To model the system, the vector of generalized coordinates is chosen as q = [ \u03b81, \u03b82, \u03c61, \u03c62 ]T , whose components are explained in the following. The angle \u03b81 accounts for the rotation of the pendulum around the y axis with respect to the zero frame reference (x0,y0,z0), while \u03b82 defines the rotation of the pendulum around the xaxis with respect to the pendulum frame reference (x1,y1,z1) as depicted in Fig. 1a. Accordingly, the rotation matrices corresponding to \u03b81 and \u03b82 can be expressed as 0 1R = c1 0 s1 0 1 0 \u2212s1 0 c1 , 1 2R = 1 0 0 0 c2 s2 0 \u2212s2 c2 , (1) where s1 = sin \u03b81, s2 = sin \u03b82, c1 = cos \u03b81, c2 = cos \u03b82, and the total transformation 0 2R = 0 1R 1 2R maps a vector from the pendulum reference frame to the zero reference frame. The angular velocity of the pendulum can be expressed in the zero reference frame as w\u0302p =0 2 R (( 1 2R )\u22121 [ 0 \u03b8\u03071 0 ]T + [ \u03b8\u03072 0 0 ]T) . (2) The kinetic energy T of the system can be found by adding the kinetic energy of the pendulum and that of the two flywheels as T = 1 2 w\u0302T p Ipw\u0302p + 1 2 \u03c6\u03071 T I1\u03c6\u03071 + 1 2 \u03c6\u03072 T I2\u03c6\u03072, (3) where w\u0302p, \u03c6\u03071, \u03c6\u03072 are the velocities of the pendulum and reaction wheels, respectively, and Ip, I1, I2 are the inertia tensors of pendulum and reaction wheels, respectively, around their centers of rotation", + " For this reason, we do not impose any terminal constraint, to maintain a relatively large domain of attraction, and verify the system stability experimentally. The nonlinear program associated to the NMPC problem (19) has a much lower complexity than the OCP (14), and can be solved in real time by using the methods described in Section III-A. CONTROLLER IMPLEMENTATIONS To implement the control algorithms and conduct experiments, a prototype of the system described in Section II was designed and built (see, Fig. 1b). The mechanical design was performed using Solidworks computer-aided design software, and the plastic parts were printed using an Ultimaker S5 3D printer. The base platform consists of universal joints, which allow the spherical motion of the pendulum about the pivot point. Four extension springs, in pre-loaded state, were mounted at the four corners of the base. To measure the angular displacements of the pendulum, two encoders were installed on the shafts of the universal joint. The inverted pendulum consists of two flywheels arranged orthogonal to each other, and attached to a rigid aluminum bar" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0003817_j.finel.2020.103494-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0003817_j.finel.2020.103494-Figure5-1.png", + "caption": "Fig. 5. Schematic view of pitch motion.", + "texts": [ + " Since FrontFlow/blue, ADVENTURE_Solid and ADVENTURE_Coupler have been developed to aim at handling large-scale problems, our analysis system can be also applied to them. Recently, we have been working on solving over one giga-degrees of freedom (DOFs) FSI problem using K computer [46]. In the future work, we will extend the present analysis to more realistic one which has a complex geometry. Although the extension engenders the increase of DOFs, it is straightforward to apply the analysis system. A flapping-wing motion can be described as a combination of three rotational degrees of freedom, namely stroke (yaw) (Fig. 4), pitch (feathering) (Fig. 5), and lead-lag (roll) (Fig. 6). According to one study [1], pitch plays a large part in the enhancement of lift generation. The pitch can be actively generated (active pitch) using actuators or passively generated (passive pitch) by the stroke if a deformable wing is adopted, as shown in Fig. 7. From a viewpoint of simplifying MAVs design, the passive pitch is very advantageous because of keeping the number of actuators low. The passive pitch has been investigated using experiments and numerical studies [15,17]", + "8 equation: CA = 1 \u2212 e\u2212 t \ud835\udf0f , (23) where \ud835\udf0f is the normal cycle time and it is set to 2.0 in the present study. During hovering flight, the upstroke and downstroke are symmetric. The upstroke (downstroke) is defined as the stroke from the negative (positive) maximum stroke angle to the positive (negative) maximum stroke angle. The quick turns from upstroke to downstroke and vice versa might cause severe distortion of the fluid mesh near the leading edge of the wing. The axis of active pitch is the leading edge of the wing, as shown in Fig. 5. The pitch angle \ud835\udf03p is defined by the following equation: \ud835\udf03p = CA\u0398p sin(2\ud835\udf0bfpt + \ud835\udefd) (deg), (24) where \u0398p is the maximum pitch angle, fp is the pitch frequency, and \ud835\udefd is the phase difference between stroke and pitch motions. When we consider a passive pitch motion, the pitch angle is generated passively according to the deformation of the wing, as shown in Fig. 7. The passive pitch angle is defined as the angle between two lines, namely a line on the rigid part of the wing chord and a normal line parallel to the y-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001160_s40194-013-0107-6-Figure5-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001160_s40194-013-0107-6-Figure5-1.png", + "caption": "Fig. 5 Movement of the cutting head and cutting directions in the experiments. Positive AOI values (\u03b2) result in a pushing cutting position in X-direction and negative to a pulling position. Bevel cutting was done in Y-direction. z-axis indicates the cutting head height adjustment", + "texts": [ + " The kerf dimensions and measured surface roughness for each case was estimated according to EN ISO 9013 standard on thermal cutting [14]. 4.2 Parameters The test materials in cutting tests were AISI 304 (1.4301) sheets and tubes. The nominal thicknesses of the sheets were 3.0 and 6.0 mm, and the measured thicknesses were 2.9 and 5.9 mm. The tube diameter was 115 mm with 2 and 6 mm nominal wall thicknesses in most of the experiments. Measured wall thicknesses for tubes were 1.9 and 5.7 mm. The material properties are shown in Table 2. Cutting directions and tilting angles are explained in Fig. 5. When cutting in X-direction, positive values of \u03b2 indicate that the cutting is performed in pushing cutting position, and negative values indicate the pulling position. In bevel cutting, the cutting direction Y is used. The cut kerf sides are referred as back side of bevel on the positive side of \u03b2 and front side with the negative \u03b2 values. The angles used in preliminary tube cutting tests were 0\u00b0, 45\u00b0, 90\u00b0, 120\u00b0, 150\u00b0, and 180\u00b0 starting from normal vertical position being 0\u00b0. The aim of the flat sheet cutting was to obtain the highest possible speed and the quality of the cut was considered of minor importance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0000639_1350650113480290-Figure1-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000639_1350650113480290-Figure1-1.png", + "caption": "Figure 1. Auto water pump bearing seal.", + "texts": [ + " In this simulation model, the topography of sealing lip and shaft surface was simulated by using autocorrelation function, the Reynolds equation and the global thermal approach was applied. A WR auto water pump bearing seal was taken as an object to calculate the lubrication parameters, such as film pressure, film thickness, friction force, heat dissipation, average temperature and leakage rate with different initial interference and different rotating speed. The relation between contact interference, heat dissipation and lubrication performance is studied. Figure 1 shows auto water pump bearing seal constituted by metal frame and rubber sealing body, the sealing zone, shaped as a belt, is formed by interference fit of sealing lip and shaft. Macroscopic structure simulation of lip seal Sealing is realized by interference contact of sealing lip and shaft, the axial structure of sealing lip is shown in Figure 2. Here, 1 expresses the angle of sealing lip at Mount Royal University on June 9, 2015pij.sagepub.comDownloaded from and shaft toward lubricant, whereas 2 expresses the angle of sealing lip and shaft toward air, zbmin is the position where the initial average film thickness is the minimum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0002483_itec-india.2015.7386862-Figure10-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0002483_itec-india.2015.7386862-Figure10-1.png", + "caption": "Figure 10. Electro-Mechanical Transmission (Renk 1100)", + "texts": [], + "surrounding_texts": [ + "The drive system of a tracked vehicle has to fulfil tasks which far exceed those known from wheeled vehicles. Apart from forward and reverse driving, it also assumes the relevant safety functions of braking and steering, and thus considerably contributes to the mobility performance characteristics of a tracked vehicle. Electromechanical drive systems have the following specific advantages with respect to mobility - Continuously adjustable driving and steering operation; Recovery of braking power; Crawling operation with the combustion engine turned off and the energy storage system installed; Conversion of the entire combustion engine output power into electrical energy; Flexibility for vehicle integration with multi-engine concepts; Integratable into an all-electric combat vehicle (AECV). The combination of mechanical and electrical components result in synergy effects which currently still have advantages over a purely electric drive with respect to safety, weight, design volume and cost. Thus, electromechanical drives for future tracked vehicles are the subject of various studies. As a consistent step in further development, the realization of an electromechanical drive system (joint development by the companies Renk, Augsburg and Magnet Motor, Starnberg) is planned. The electromechanical drive system combines the benefits of proven mechanical drive technology with those of the future oriented electrical drive technology. The specific advantages of this concept are: \u2022 The two-speed transmission provides for two operating modes. The first mode will only be required in extreme situations, such as difficult cross-country terrain. The second mode with direct transmission can be used in paved light terrain and on primary roads (highways). \u2022 Maximum speed is continuously adjustable. \u2022 The requirements for electrical components can be reduced. \u2022 In both modes the MEP main engine can be operated at optimum efficiency. \u2022 Independent drive, steering and braking systems provide for a high level of system safety. \u2022 Stabilized straight-ahead movement. \u2022 Maintenance of the regenerative mechanical steering principle with neutral shaft. \u2022 Flexible concept design obtained by integrating all components in one block (\u201cpower pack\u201d) or independent arrangement by the separation of combustion engine(s) and drive block." + ] + }, + { + "image_filename": "designv11_34_0000318_j.finel.2015.07.008-Figure2-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0000318_j.finel.2015.07.008-Figure2-1.png", + "caption": "Fig. 2. The dynamic model of the five shaft geared rotor system.", + "texts": [ + " 1, which is a prototype of an integrally geared centrifugal compressor with 5 shafts coupled by helical gears in parallel arrangement. Each shaft is supported by two TPJBs. The test rig is driven by an AC motor connected to the input shaft (I). The dummy impellers or discs are fixed on the three output shafts (O1, O2, O3). These shafts are all connected to a big main gear (M) by different gears. The transmission ratios of the three output shafts are 2.72 (I to O1), 3.25 (I to O2) and 5.10 (I to O3). The dynamic model of the five shaft geared rotor system is shown in Fig. 2. kiL, ciL and kiR, ciR are the stiffness parameters and damping parameters of the left TPJBs and right TPJBs on the five shafts respectively, which are dependent on the loading characteristic. Fi is the meshing force. Unlike the traditional one-to-one gear pair's transmission, the geared system in Fig. 2 is consisted of a bull-gear (M) and four pinion gears, and the bull-gear meshes with the other four gears at the same time, which cause a much stronger coupling effect. This compact structure can offer several advantages such as smaller floor space, lower energy consumption, and higher power density. However, it also means that much more critical speeds must disperse in a certain interval of speed, which brings much more difficulties in critical speed designing. The finite element method is used to set up the dynamic model of the geared rotor system shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_34_0001946_ecce.2014.6953971-Figure3-1.png", + "original_path": "designv11-34/openalex_figure/designv11_34_0001946_ecce.2014.6953971-Figure3-1.png", + "caption": "Fig. 3 discretization of IPMSM.", + "texts": [ + " In field weakening operation, we cannot assume a sinusoidal waveform of the flux density so we will use models taking into account the influence of the flux density derivative on iron losses as LS and generalized Bertotti models. Although the LS model gives a good description of the magnetic losses, the generalized Bertotti model could be more suitable because it is faster and easier to calibrate with the material type. The developed model for our application obtains the FDW from a nonlinear nodal network with the methodology described in [11] and solving the problem in magneto-static. The IPMSM of Fig. 1 can be discretized in different areas as described on Fig. 3. These areas represent the main flux ways as described in Fig. 2 which have been defined from a preliminary study of the IPMSM with finite elements software (FLUX 2d). The nodal network has been defined from a first study with FE approach is given in Fig. 4. So this model cannot provide the temporal FDW in the different parts of the stator but the spatial FDW. However it is possible to assume that the temporal FDW in the teeth and the stack are identical with the spatial FDW. With this assumption, only the slot harmonics are neglected" + ], + "surrounding_texts": [] + } +] \ No newline at end of file