diff --git "a/designv11-92.json" "b/designv11-92.json" new file mode 100644--- /dev/null +++ "b/designv11-92.json" @@ -0,0 +1,8208 @@ +[ + { + "image_filename": "designv11_92_0001073_epepemc.2008.4635363-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001073_epepemc.2008.4635363-Figure1-1.png", + "caption": "Fig. 1. Hypothetical SRM2-2.", + "texts": [ + " The suggested approach helps to avoid the sophisticated mathematical description, provoked by the fact that the stator inductance depends on both, position of the rotor and the stator current value [2, 3]. The last dependence leads to partial differential equation description, which sometimes is rather complex task, even for contemporary simulation facilities. Furthermore, neural network description helps to reflect the nonlinearity character of the magnetic circuits, as well as the influence of the air gap. II. THE APPROACH ESSENCE A. Hypothetical nonlinear neural network motor model The starting point of this approach is the hypothetical motor model illustrated in Fig. 1. In fact it is a SRM2-2 motor that has 2 stator and 2 rotor poles. This model is not of practical use but it is a base point for further development of different SRM models. The mathematical description of this hypothetical single-phase motor is given below (system of equations (1)). It is a result of a previous research of the authors [1]. )().(),( NILIL (1-a) )()( NetNN (1-b) )(' )( NetN d dN (1-c) )()( INetLIL (1-d) )( )( ILNet dI IdL (1-e) dt d (1-f) dt d (1-g) dI IdL NIIL d dN ILR I U I dt dI )( )", + " The structure of the acceleration block is depicted in Fig. 5. It is seen from it that tribo (friction and bearings) effects are also taking into account. The complete block diagram of the hypothetical motor which is described by the whole system \u2013 from (1-a) to (1-i), can be seen in Fig. 6. If there is a need to describe the hypothetical motor shown in Fig. 7, then the suggested convenient description will be like (2): ), 2 (),(),( 2 2 2 2 2 4 UMUMUM S R S R S R (2) The last model is obtained as a superposition of the model that is pictured in Fig. 1, taken two times and with phase difference of superscript and the subscript ( 2 4 S RM ) the motor is a single phased SRM2-4 motor with 2 stator poles and 4 rotor poles. The simplified model can be seen in Fig. 8. The position feedback gives opportunity for 4 stable aligned 788 2008 13th International Power Electronics and Motion Control Conference (EPE-PEMC 2008) rotor locations. Evidently this description is very close to a single phased SRM6-4 motor type. What happens if the motor consists of three phases like the one shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003942_iccairo.2018.00054-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003942_iccairo.2018.00054-Figure2-1.png", + "caption": "Fig. 2. MESSENGER spacecraft and MLA on the main instrument deck.", + "texts": [ + " The intense heat from the Sun requires the spacecraft to point its sunshade toward the Sun at all times; during noon\u2013midnight orbits, this requirement constrains the payload deck to point within \u223c10\u25e6 about the ecliptic south pole, with MLA ranging at a slant angle as high as 70\u25e6. The MLA measures the topography of Mercury via laser pulse time-of-flight and spacecraft orbit position data. The primary science measurement objectives for MLA are to provide a high-precision topographic map of the northern polar regions, to measure the long-wavelength topographic features of the mid-to-low latitude regions, and to detect and quantify the planet\u2019s forced librations. Fig. 1 shows a photograph and a mechanical drawing of MLA. Fig. 2 shows U.S. Government work not protected by U.S. copyright. the MESSENGER payload instruments and the location of MLA on the spacecraft. MLA operates under a harsh and highly dynamic thermal environment, due to the large variation in heat flux from the Mercury surface from daytime to nighttime and from deep space background. The transmitter and receiver optics undergo rapid and uneven swings in temperature during science measurement (tens of degrees per hour at the laser-beam expander and the receiver telescope)", + " The calibration tests can be categorized into several groups: pointing and timing references; range biases; laser characteristics; and receiver characteristics. A series of calibration tests was conducted, and some were repeated at different temperatures with the instrument in the vacuum chamber. The calibration results are described here. The MLA coordinates, with respect to the spacecraft coordinate system, are obtained from a combination of the MLA measurements and spacecraft survey results after payload integration. The definitions of the MESSENGER spacecraft coordinate system are labeled in Fig. 2. The XY plane is the same as that defined by the launch vehicle separation plane, and the origin is at the center of the adapter ring. MLA orientation in the spacecraft coordinate system is shown in Fig. 1. All the MLA physical dimensions and angular directions can be referenced to the alignment reference cube attached to the side of the main housing, as shown in Fig. 1. The coordinates of the MLA alignment reference cube in the MESSENGER spacecraft coordinate system have been measured to be [\u221215" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003903_ecc.2019.8795860-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003903_ecc.2019.8795860-Figure1-1.png", + "caption": "Fig. 1. Quadrotor Coordinate System", + "texts": [ + "1: [21] If there exists a continuous function Q : R+ \u00d7 Rn \u2192 R (V, x)\u2192 Q(V, x) satisfying the conditions C1) Q is continuously differentiable outside the origin; C2) for any x \u2208 Rn\\{0} there exists V \u2208 R+ such that Q(V, x) = 0 C3) let \u2126 = {(V, x) \u2208 R+ \u00d7 Rn : Q(V, x) = 0} and lim x\u21920,(V,x)\u2208\u2126 V = 0+, lim V\u21920+,(V,x)\u2208\u2126 \u2016 x \u2016= 0, lim \u2016x\u2016\u2192\u221e,(V,x)\u2208\u2126 V = +\u221e C4) \u2202Q(V,x) \u2202V < 0 for all V \u2208 R+ and x \u2208 Rn\\{0}; C5) There exist c > 0 and 0 < \u00b5 \u2264 1 such that sup t\u2208R+ \u2202Q(V, x) \u2202x f(t, x) \u2264 cV 1\u2212\u00b5 \u2202Q(V, x) \u2202V for all (V, x) \u2208 \u2126 then the origin of the system is globally uniformly finite-time stable and T(x0) \u2264 V \u00b50 c\u00b5 , where Q(V, x) = 0. III. PROBLEM STATEMENT The dynamic model of quadrotor is well established in the literature (see, e.g. [24], [25]). Since the hub forces and moments are very small compared to actuator moments and forces, they are negligible in the considered model [26]. Two different coordinate systems are used in the system model: the initial and body coordinate systems, denoted by i and b respectively (see Fig.1). The position and the attitude in initial frame of quadrotor are given by (x, y, z) and (\u03c6, \u03b8, \u03c8)(roll-pitch-yaw) respectively. The total thrust F of quadrotor in body frame is provided by four thrusts f1, f2, f3, f4 of propeller, F = f1 +f2 +f3 + f4. The thrust of each propeller is fi = k\u03c92 i , where k is the thrust coefficient and \u03c9i is the rotation speed of propeller. The force in the inertial frame is given by R(\u03c6\u03b8\u03c8)F , where R(\u03c6, \u03b8, \u03c8) = C\u03b8C\u03c8 \u2212S\u03c6S\u03b8C\u03c8 + C\u03c6S\u03c8 \u2212C\u03c6S\u03b8C\u03c8 \u2212 S\u03c6S\u03c8 \u2212C\u03b8S\u03c8 S\u03c6S\u03b8S\u03c8 + C\u03c6C\u03c8 C\u03c6S\u03b8S\u03c8 \u2212 S\u03c6C\u03c8 S\u03b8 S\u03c6C\u03b8 C\u03c6C\u03b8 (2) with the notations C\u03b8 = cos \u03b8, S\u03b8 = sin \u03b8, etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003394_6.2019-1747-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003394_6.2019-1747-Figure1-1.png", + "caption": "Figure 1: Example of a closed-chain structure with tape spring hinge integration.", + "texts": [ + "5, 6 A central challenge for this concept is the deployment dynamics and deployment actuation of the folded structure and spacecraft system. A novel lightweight solution is to integrate strain energy hinges to facilitate folding and actuate the deployment.7 High strain composite tape spring hinges are an intriguing innovation in hinge technology for deployable space structures. Compared to standard piano hinges, these hinges are lightweight, eliminate rotational mechanical contact surfaces, and are self-actuating. A simple example of how this concept could be implemented physically using the miura-ori pattern is illustrated in Figure 1, and it is noted that even with minimal hinge actuation, 10 hinges are used to actuate the 12 panel assembly. Of additional note is that the thickness of the fold panels requires some minimal gap thickness between each panel to facilitate folding, enabling small displacements in multiple degrees of freedom. Deployment dynamics of such a system would typically be studied through finite element analysis (FEA). However, for a structure with multiple high strain composite hinges, FEA modeling would require significant computational time and skill" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002866_s1028335818100099-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002866_s1028335818100099-Figure1-1.png", + "caption": "Fig. 1. Schematic model of a spherical robot on an inclined plane.", + "texts": [ + " In [1], equations of controlled motion of a spherical robot of pendulum type on an inclined plane are presented and integrals of motion and partial solutions (with constant control) are found. A linear stability analysis of partial solutions is carried out.In this paper, using the results obtained in [1\u20134], an algorithm is presented for constructing an explicit control during motion along a given trajectory on an inclined plane, in particular, in a straight line and a circle. Consider the motion of a spherical robot of pendulum type rolling without slipping on an inclined absolutely rough plane. The spherical robot is a spherical shell of radius (Fig. 1) with an axisymmetric pendulum fixed at its center. The pendulum (Lagrange top) has mass m, the distance from the center of the spherical shell to the center of mass of the pendulum is , and in the system of principal axes of the top the central tensor of inertia has the form . The system is set in motion by forced oscillations of the pendulum. In this paper, we address a model problem in which it is assumed that the mechanism setting 1 The article was translated by the authors. oR tR = , , +0 diag( )i i i ji 43 aV", + " Steklov Mathematical Institute RAS, Moscow, 119991 Russia * e-mail: archive@rcd.ru the pendulum in motion produces control torque Q, applied at the point C of attachment of the pendulum. All vectors describing the motion of the spherical robot will be referred to a fixed (inertial) coordinate system (the -plane coincides with the inclined plane of rolling of the spherical robot, the unit vector is directed to the observer perpendicularly to the plane of the figure, and the unit vector is directed along the outer normal to the inclined plane), see Fig. 1. The velocity of the center of mass of the top is determined from the condition that the velocity of the center of the shell is equal to that of the point of suspension of the pendulum. This condition is expressed by the (holonomic) constraint (1) Oxyz Oxy \u03b2 \u03b3 v V = + \u00d7 ,tRv V \u03c9 n 5 where n is the vector directed along the symmetry axis of the pendulum and is the angular velocity of the pendulum. The velocity of the center of the shell is related to the angular velocity of the shell, , by the condition that there is no slipping at the point of contact : (2) where are the coordinates of the center of the sphere, , and is the angular velocity of the shell", + " (5) into (6) taking the assumption , into account, we obtain explicit dependences . Here we assume that the initial angular velocity of the shell is matched with the trajectory and that Eqs. (6) are satisfied at the initial instant of time. (2) Introduce new variables related to the variables by (7) The variables , , and have the obvious physical meaning: is the angular velocity of rotation of the pendulum about its symmetry axis, is the angle of deviation of the pendulum from the line drawn from the center of the sphere to the point of contact of the plane (see Fig. 1), and is the angle between the axis and the projection of the vector onto the inclined plane. The variables , are related to and by (8) (3) Eliminating the control from the system, add the first two equations of (3). Substitute the explicit dependences obtained earlier into the resulting equation and the third equation of (3). The resulting equation, written in the variables X, and Eqs. (8) form a closed system of five differential equations explicitly depending on time: (9) where , are, respectively, the matrix and the vector of the coefficients of the equations, which depend explicitly on time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003033_icems.2018.8549531-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003033_icems.2018.8549531-Figure1-1.png", + "caption": "Fig. 1. Cross section of test SRM.", + "texts": [ + " The efficiency and the output of switched reluctance motors are largely influenced by excitation timing. So far, the effect to the motor efficiency and losses by variable excitation interval single pulse control has been studied [1]. Several research reports have improved SRM characteristics by current mode. [2]-[7]. These paper clarified that the maximum output is improved by continuous current mode. In this paper, three phase critical current mode is proposed and the availability is verified by simulation on motor efficiency and power factor. II. SPECIFICATIONS OF TEST SRM Fig. 1 shows the cross section of test SRM and Table 1 shows the specifications. The test SRM is the double salient pole structure of stator 6 poles and rotor 4 poles. The winding wire which made 3 parallel conducting wire of 0.75 mm in diameter is coiled 20 T around a stator pole. The winding wire of pole a and pole a is connected in series. The rated output is 180 W and the rated voltage is 24 V. The origin position of the motor makes unaligned position 0\u00b0(mech.). An overlap angle is 13.66\u00b0(mech.). The direction of rotation defines clockwise rotation as positive rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure5-1.png", + "caption": "FIGURE 5 Gear tooth involute profile measurement principle", + "texts": [ + " ISO 1328 [1] and ISO 10064 [2] give gear tooth profile and gear tooth helix measurement procedures. ISO 1328 defines the profile deviation as the amount by which the measured profile deviates from the design profile and defines the helix deviation as the amount by which the measured helix deviates from the design helix. The profile deviation is measured normal to the actual profile in the transverse plane in the profile evaluation range, whereas, the helix deviation is measured normal to the actual helix in the direction of the transverse base tangent in the helix evaluation range. Figure 5 shows the profile measurement details in relation with the gear geometry. The mean profile and the mean helix represent the shapes of the design profile and the design helix respectively and are aligned with the measured trace. Form deviation, slope deviation, and total deviation are used to characterize the measured trace. 1\u00a0:\u00a0Design profile 2\u00a0:\u00a0Actual profile 3\u00a0:\u00a0Mean profile 1a\u00a0:\u00a0Design profile trace 2a\u00a0:\u00a0Actual profile trace 3a\u00a0:\u00a0Mean profile trace 4\u00a0:\u00a0Origin of involute 5\u00a0:\u00a0Tip\u00a0point 5\u00a0\u2212\u00a06\u00a0:\u00a0Usable profile 5\u00a0\u2212\u00a07\u00a0:\u00a0Active profile C\u00a0\u2212\u00a0Q\u00a0:\u00a0Base tangent length to point\u00a0C \u03beC\u00a0:\u00a0Involute roll angle to point\u00a0C Q\u00a0:\u00a0Start of roll angle A\u00a0:\u00a0Tooth\u00a0tip\u00a0or start of chamfer C\u00a0:\u00a0Reference point E\u00a0:\u00a0Start of active profile F\u00a0:\u00a0Start of usable profile LAF\u00a0:\u00a0Usable length LAE\u00a0:\u00a0Active length L\u03b1\u00a0:\u00a0Evaluation range LE\u00a0:\u00a0Base tangent lenth to start of active profile F\u03b1\u00a0:\u00a0Total profile deviation ff\u03b1\u00a0:\u00a0Profile form deviation fh\u03b1\u00a0:\u00a0Profile slope deviation Figure 6 shows an involute tooth profile with a vertical shift after assembly overlaid on an involute tooth profile of an unassembled gear for comparison" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000983_icelmach.2008.4800111-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000983_icelmach.2008.4800111-Figure2-1.png", + "caption": "Fig. 2. Examples of rotor cross-sections for: a) Surface Mounted PMSG, b) Interior PMSG.", + "texts": [ + " In the case of standalone generators, for which the magnetizing reactances in direct d and quadrature q rotor magnetic axes are equal: Xmd = Xmq or very similar, the increase of load current for load type R or R-L causes the decrease of output voltage. The load characteristic of output voltage vs. output power VLL = f(P2) is nonlinear and it is falling with the increase of output power (Fig. 1). This falling shape of VLL = f(P2) characteristic is common for the PMSGs with permanent magnets mounted on the surface of cylindrical rotor core (Surface Mounted PMSG, Fig. 2.a). In the case of these PMSGs in which magnets aren\u2019t mounted on the surface of cylindrical rotor but they are mounted in the dedicated slots placed inside of solid or laminated rotor core (Interior PMSG, Fig. 2.b), the dependence between magnetizing reactances in d and q rotor axes has usually a form: Xmd < Xmq. Due to this dependence, in such generators the armature reaction is significantly stronger in the q-axis and it is known, that strong q-axis armature reaction can cause the increase of output voltage at stand-alone generator terminals with increase of the load current. So, for the standalone Interior PMSG connected to R or R-L type load, the VLL = f(P2) characteristic is typically more \u201cflat\u201d comparing to the same characteristic of Surface Mounted PMSG with similar overall dimensions and similar stator construction (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002994_fpmc2018-8876-Figure14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002994_fpmc2018-8876-Figure14-1.png", + "caption": "FIGURE 14. ELLIPTICAL PROFILES PUBLISHED IN LITERATURE", + "texts": [ + " The OFs and constraints were also normalized to the maximum allowable values to help the GA find the feasible design space more quickly and a more uniform distribution of designs by the crowding distance criterion [21]. The results of the optimization are also compared to two pump geometries found in literature by Jung et al. that are claimed to out-perform two circular-toothed geometries used in automotive applications based on size, flow ripple, and relative sliding [26]. The profiles are shown in Fig. 14 and were the only published elliptical profiles with dimensions the authors were able to find. Elliptical-toothed profiles may already be used in industrial applications, but their specific geometries are likely guarded as trade secrets. 8 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 03/05/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use Of the 500,000 evaluated designs in the optimization, 28% were feasible, and 13% were non-dominated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002052_iciea.2018.8398116-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002052_iciea.2018.8398116-Figure2-1.png", + "caption": "Fig. 2. Transformation in natural coordinate system and rotating coordinate system", + "texts": [ + " The reference values of the five phase currents of the PMSM are obtained by the clarke transformation. Then the difference between the reference currents and the actual current is input to the hysteresis comparator. Finally, the PWM signals are generated to control the inverter, thus controlling the five phase PMSM [12]-[13]. Fig. 1. Block diagram of control strategy of five phase PMSM drive system. For implementing the control strategy of five phase PMSM, it is crucial important to consider clarke transforms, which is different from that used in three-phase PMSM. Fig.2 shows that the transformation in natural coordinate system and stationary coordinate system. 2018 13th IEEE Conference on Industrial Electronics and Applications (ICIEA) 2427 The clarke transform between the natural coordinate system and the rotating coordinate system is expressed as \u2212\u2212\u2212 \u2212\u2212\u2212 \u2212\u2212\u2212 \u2212\u2212\u2212 \u2212 = q d e d c b a i i i i i i i )4sin()4cos( )3sin()3cos( )2sin()2cos( )sin()cos( sincos 5 2 \u03b1\u03b8\u03b1\u03b8 \u03b1\u03b8\u03b1\u03b8 \u03b1\u03b8\u03b1\u03b8 \u03b1\u03b8\u03b1\u03b8 \u03b8\u03b8 (14) where 5/2\u03c0\u03b1 = , id and iq are the currents in the rotating coordinate system V" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001726_memsys.2018.8346783-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001726_memsys.2018.8346783-Figure5-1.png", + "caption": "Figure 5: Cross-section of the entire assembly consisting of the sensing chip, the holder with the electric and fluidic connections, the excitation LED, a filter and the camera", + "texts": [ + " The optical setup consists of an excitation LED with a center wavelength of 395 nm, a phosphorescent film, an optical filter with a cut-off wavelength of 600 nm and a Raspberry Pi camera to detect the emitted light of the film. The chip is clamped into a 3D-printed holder through which the sample liquid is inserted. This connection is sealed with O-rings. Furthermore, the holder contains four spring-probes to contact the chip electrically. The optical components are mounted on an additional 3D-printed holder as shown in fig. 5. The read-out of the temperature sensor is realized with a Wheatstone-bridge and AD-converter. A PI-control to set the temperature is embedded in the Raspberry Pi which also controls the current through the excitation LED and takes the photos. (the assembly is symmetrical regarding the cross-section surface) The image of the film is recorded with the Raspberry Pi and the mean value of each spot is analyzed. Fig. 6 shows the average of 5 recorded spot intensities for different temperatures and oxygen concentrations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003864_012022-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003864_012022-Figure2-1.png", + "caption": "Figure 2. The non-central collision on billiard games with \u03b8 (angle of cue ball cuff ball towards cue ball direction before collision) and \u00d8 (object ball scattering angle to cue ball direction before collision)", + "texts": [], + "surrounding_texts": [ + "Billiard games is a sport that falls into the category of sports concentration. This sports branch is played on the table and with special auxiliary equipment and its own rules. This game is divided into various types such as Carom, English Billiard and Pool. Mass of billiard sticks is 28.35 grams and has a length of 147-152 cm. The billiard table has a rectangular shape with a length of 270 cm and a width of 135 cm. Billiard balls have a massive ball shape with a diameter of 5.7 cm. Mass of billiard is not the same that is mass of cue ball is 0.17 kg and mass of object ball is 0.16 kg [12]." + ] + }, + { + "image_filename": "designv11_92_0003890_02533839.2019.1598290-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003890_02533839.2019.1598290-Figure8-1.png", + "caption": "Figure 8. The vertical C-type machine center.", + "texts": [ + " Similar to Equation (34), the following equation can be obtained as: xf\u00fe1 \u00bc \u0393\u00fei 1Oi 1: (36) By representing Equation (34) and (36) using the state space, the following equation can be formulated: xf\u00fe1 Yi \u00bc A A C D xf Ui (37) Using the least square method, one can compute the (A, B, C, D) which is the subspace model using the N4SID approach. In next section, the SIM method is applied to identify the feed drive system. To verify the proposed method, a vertical C-type machine tool (P series) with three linear axes made by Victor in Taiwan is used as the experimental platform as shown in Figure 8. Taking the X axis as an example, the feed drive system consists of motor, coupler, ball-screw, linear guide, and moving platform. The CNC controller used is FANUC 0-iMF with the AICC-II function. The interpolation process of the FANUC controller is almost the same as that described in Section II. To validate the interpolation process, the Servo Guide software provided by FANUC is utilized to record the position commands. The maximum acceleration (Amax) and jerk time (T2) are chosen as the designed variables to generate different trajectories which can be used to test the SIM model\u2019s accuracy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000741_apccas.2008.4746131-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000741_apccas.2008.4746131-Figure3-1.png", + "caption": "Figure. 3 Screw coordinates of the first limb", + "texts": [ + " The foundation of screw coordinates In the study of kinematic analysis of mechanisms, the screw theory has been widely used for describing rigid-body motion in three dimension. A screw is a line to which attached a scale pitch parameter and usually expressed as: $ [ , , ; , , ]L M N P Q R= In which L, M, N are the direction ratios of the line and P, Q, R are the direction ratios of the result moment of the line with respect to the origin. According to the screw coordinates definition, the screw coordinates of axis line of all joints in the first limb can be computed in the fixed coordinate system O-x1y1z1 shown in Fig.3. 1$ [0, 1, 0 ; 0, 0, R]= (8) 2 11 11 11 11$ = [0 , 1 , 0 ; -L sin( ), 0 , R+L cos( )]\u03b8 \u03b8 (9) 3 11 12 11 12 11 12 11 12 12$ [sin( + ) , 0,-cos( + ); 0,Rcos( + )+L cos( )+L ,0]\u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8= (10) 4 11 12 11 12 13 13 11 12 11 12 11 12 12 13 13 13 13 11 12 $ [sin( + ) ,0,-cos( + );-L sin( )cos( + ), cos( + )+L cos( )+L +L cos( ) , -L sin( )sin( + )] \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = (11) 5 11 11 12 11 12 12 13 11 12 11 11 12 11 12 13 13 11 12 $ [0 ,1,0;-L sin( )-L sin( + )-L cos( )sin( + ), 0,R+L cos( )+L cos( + )+L cos( )cos( + )] \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = (12) Similarly, the screw coordinates of all joints in the second limb and the third limb can be calculated in the same way mentioned above" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000881_pi-b-2.1962.0229-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000881_pi-b-2.1962.0229-Figure9-1.png", + "caption": "Fig. 9.\u2014Derivation of bent-sleeve aerial. (a) Bent unipole.", + "texts": [ + " less than 2 : 1 relative to 50 Q., a frequency range of one octave is possible with this type of unipole, whereas the simple unipole will cover only a 20 % range. In practice, the gap between the two sections of the aerial in Fig. 7(c) is covered with an insulator. The loads on this are only a quarter of those at the base of the aerial if the sections are equal in length. Another form of the sleeve aerial is based on a unipole in which the upper half is bent parallel to the ground plane, as in Fig. 9(a). The radiation resistance of the quarter-wave unipole is of the order of 5Q. only. By adding an eighth-wavelength sleeve, the impedance can be brought up to the order of 50 Q. Fig. 9{b) shows a typical bent-sleeve aerial in which the sleeve has been shaped to give the best aerodynamic performance. (b) Streamlined bent-sleeve unipole. Even with the compromise which this requires, the v.s.w.r. is less than 2 relative to 50 Q over the frequency range 200- 400 Mc/s. One characteristic of sleeve aerials is that the impedance at the half-wave resonance remains low; experimentally, a bent-sleeve aerial was made to cover the range 350-1400 Mc/s with a v.s.w.r. of less than 2 : 1 . The radiation patterns were not examined over this range, and may not have been sufficiently omnidirectional to allow full use of this development" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001479_978-3-319-72730-1_26-Figure26.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001479_978-3-319-72730-1_26-Figure26.2-1.png", + "caption": "Fig. 26.2 Chord- and tangent-slope inductances", + "texts": [ + " From the equations above, a rather practical mathematical flux model can be derived. For steady state and no load, the main flux can be written using the chord-slope inductance \u03c8 f mo(I f o) = L f mo(I f o) \u00b7 I f o. As a consequence, for a loaded machine, \u03a8 f mo(I f o, Iao) = L f mo(I f o) \u00b7 I f o \u2212 w f k f (I f o) \u00b7 I 2ao = L f mo(I f o) \u00b7 I f o \u2212 K f (I f o) \u00b7 I 2ao (26.22) Small variations around an equilibrium point, on the other hand, are described using incremental or tangent-slope inductances (see also Fig. 26.2) \u0394\u03a8 f m = L f mt (I f o) \u00b7 \u0394I f \u2212 2K f (I f o) \u00b7 Iao \u00b7 \u0394Ia = L f mt (I f o) \u00b7 \u0394I f \u2212 M f a(I f o, Iao) \u00b7 \u0394Ia (26.23) The inductance L f mt is the tangent-slope inductance of the no-load characteristic for the excitation current I f o. Note that the dynamic (or incremental) mutual inductance M f a can be positive or negative, dependent on the sign of the armature current: M f a(I f o, Iao) = 2K f \u00b7 Iao = 2K f \u00b7 |Iao| \u00b7 sign(Iao). For the total excitation flux, the leakage flux needs to be added (the leakage flux can be assumed to be unsaturated)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure9-1.png", + "caption": "FIG. 9: (A) 90\u00b0 TORUS, (B) 4 TRIM PLANES, (C) 4 PLANAR BUILD REGIONS.", + "texts": [ + "asme.org on 02/13/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use These examples illustrate building well defined features. However, geometry can be segmented to be built on multiple planes to eliminate the need for support material. The need for support material for an overhang condition varies based on the process-machine-material set. This needs to be considered when developing a segmentation strategy. The segmented geometry cannot have open faces. The 90\u00b0 torus illustrated in Fig. 9 (a) is segmented into 4 sections (Fig. 9 (b)), and a build plane established for each section. The build planes for the torus are the XY plane (0\u00b0), and rotated 22.5\u00b0, 45\u00b0, and 67.5\u00b0 around the Z axis. The slices for each segment in Fig. 9 (c) are parallel to each build plane. The amount of support material for this 90\u00b0 torus would be minimal if it were built \u2018laying down\u2019 or the top view orientation, but if the torus were a thin walled component, or had some internal geometry, support material would be required in any build orientation, and it would be challenging to remove. Employing a rotary positioning system, and including side tilt, provides a fabrication solution for the geometry shown in Fig. 10 (b) while using conventional slicing and tool path options" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003352_6.2019-0776-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003352_6.2019-0776-Figure2-1.png", + "caption": "Fig. 2 Example casing CADmodel with cross-sectional view. \"Top\" and \"base\" nomenclature is derived from DMLS machine build direction.", + "texts": [ + " Functional tests captured the cylinders\u2019 overall dynamic response to explosive loading, DIC testing provided detailed images of pre-rupture casing expansion needed for DIC analysis, and fragment tracking tests recorded post-rupture fragment movement for velocity calculations using MATLAB-based algorithms. All experiments were conducted within a custom-built Lexan container positioned inside of an indoor explosives chamber for optimal optical and lighting conditions. Due to software and time constraints, DIC analysis was not completed in time for this publication. A. Sample Design & Construction As illustrated in Figure 2 below, cylindrical casings were first designed as CAD models in Autodesk Inventor before conversion into STEP files for DMLS machine compatibility. Small-scale test samples, less than 0.5 grams, were used to maximize data collection rate during the allocated test window. Previous work conducted at Sandia National Laboratories (SNL) showed fragmentation at these scales displays many qualitative similarities to larger scale experiments. Their experiments utilize modified Teledyne RISI RP-80 exploding bridgewire detonators, depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001510_978-981-10-5768-7_13-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001510_978-981-10-5768-7_13-Figure1-1.png", + "caption": "Fig. 1 Overview of the test rig. 1 foundation; 2 test bearing; 3 test frame; 4 drive system; 5 static load equipment; 6 rotor and support; 7 excitiers", + "texts": [ + " Therefore, the purpose of the present paper is to investigate experimentally the validity of the model described in Khatri [13] by testing practical journal bearings and comparing the performance of LST bearings with that of the theoretical predictions and with the performance of standard non-textured bearings. Experimental results have been compared with a theoretical model. The experimental results are in good agreement with the theoretical model under relevant operating conditions. The same bearing geometrical parameters and test rig studied by Jiexi Shen [14] are used here for the experiment. Figure 1 shows the 3D model of the journal bearing test rig. The test rig is mainly composed of the following parts: 1. Two pedestals to support the test rotor through two same ball bearings that are lubricated through the oil supply system. 2. An electromotor used to drive the rotor. The maximum speed is 3000 rpm. 3. A pneumatic loader to apply a static load to the stator in the positive y direction. The maximum available load is 5 kN. 4. Dynamic load system, by 90\u00b0 two degrees of the vibration exciter, the use of non-contact exciter, the machine of the vibration exciter is JZQ-70, and the maximum output exciting force is 700 N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002330_978-981-13-0341-8_8-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002330_978-981-13-0341-8_8-Figure6-1.png", + "caption": "Fig. 6 Laboratory experimental set up to collect AE data [12]", + "texts": [ + " The real values are first fed into the sigmoid function \u03c3 (x) that squashes the real values output to [0, 1] and the model prediction becomes y\u0302(x) \u03c3 (WT x + b). Then, the cross-entropy loss is defined by the sigmoid function as L\u03b8 (y, y\u0302) 1 M M\u2211 m 1 [\u2212y(m) log y\u0302(y(m)) \u2212 (1 \u2212 y(m)) log(1 \u2212 y\u0302(y(m))) ] (9) Finally, the DNN is trained to optimize entire weight and bias using the backpropagation algorithm. To evaluate the proposed method, data is collected from the test rig of the laboratory, as shown in Fig. 6. A three-phase motor is coupled with the drive-end shaft of the gearbox. The gearbox transmits torque to the non-drive-end shaft. The cylindrical roller element bearing is installed at both ends of each shaft. The non-drive-end shaft is connected with a belt and pulley to rotate the blade as the loaded condition of a bearing. The displacement transducer is placed on the non-drive-end shaft tomeasure the rotating speed of the bearing. An acoustic emission (AE) sensor is installed on the bearing housing to record AE signals stating bearing health condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure17-1.png", + "caption": "FIG. 17: UNROLLED GEOMETRY, TRIMMED OFFSET TOOL PATH SEGMENTS, AND ROLLED TOOL PATH SEGMENTS", + "texts": [ + " Ideally, the AM rotary tool paths could support nonrotational features. One solution for set for the examples presented here can be achieved through unroll-roll functions for the feature edge geometry. Consider the flange in Fig. 16. When unrolling the bounding swept geometry (a circle), a straight line results; whereas, the unrolled edge geometry for the square flange appears as a set of semi-circles, where the farthest protrusion points for the flange touch the straight line. The unroll operation is performed for the cam example (Fig. 17). Offsetting the upper bound by the desired slicing layer height, and trimming these offset curves to the unrolled geometry edges will generate the stop-start points for the deposition segments for each deposition layer. Using a crude offset distance, this is illustrated for the cam profile. Attention must be paid to managing the unrolled geometry initial positions, as this links the angular relationships to the deposition start-stop points for a given roll diameter. The unrolled geometry for a 3-lobed cam is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003308_6.2019-1073-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003308_6.2019-1073-Figure1-1.png", + "caption": "Figure 1. Complete Geometry and Simplified Geometry of the UAV.", + "texts": [ + " U D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 10 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 10 73 AIAA Scitech 2019 Forum 7-11 January 2019, San Diego, California 10.2514/6.2019-1073 Copyright \u00a9 2019 by Design, Analysis and Research Corporation. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission. AIAA SciTech Forum 2 American Institute of Aeronautics and Astronautics A polyhedral mesh is applied to the rotor and the air volume surrounding the rotor as shown in Figure 1. With Blade Element Momentum (BEM) in STAR-CCM+, a virtual disk model is used to represent the rotor. The approach starts with a steady state, coarse mesh and static propellers setting. After establishing a desired mesh set up, a finer mesh is applied. Next, the rotating propellers at 5,300 rpm are introduced in the fine mesh setting with the unsteady Reynolds Average Navier Stokes solver. This rpm setting represents the hover condition of the DJI Phantom II. Finally, a full turbulence solver model in detached eddy simulation is activated to run util a solution convergence is obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000415_1.1718791-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000415_1.1718791-Figure2-1.png", + "caption": "FIG. 2. Close-up view of electrical contact area.", + "texts": [], + "surrounding_texts": [ + "device were about 0.017 in. in diam; smaller samples would require a narrower tunnel and smaller vent dimensions.\nWe regularly obtain full half-widths of about 0.35 Oe without further treatment. Annealing at 900\u00b0C for 1 h with a slow cool may reduce this to about 0.3 Oe, which we feel is close to the intrinsic linewidth for the purity of materials we have.\nWe wish to thank John Clark and Gordon Harrison of the Sperry Microwave Electronics Division in Clearwater, Florida, for providing us with some of the YIG crystals and Norman Goldberg and William McCann for others. We acknowledge with thanks the assistance of James Douglas in all phases of the work.\n* Present address: U. S. Naval Ordnance Laboratory, White Oak, Sliver Spring, Maryland.\n1 R. C. LeGraw, E. G. Spencer, and C. S. Porter, Phys. Rev. 110, 1311 (1958).\n2 W. L. Bond, Rev. Sci. lnstr. 22, 344 (1951). 3 P. Senio and C. W. Tucker, Rev. Sci. lnstr. 24, 549 (1953). 4 W. L. Bond, Rev. Sci. lnstr. 25, 401 (1954). \u2022 F. Reggia and W. Stadler, Rev. Sci. lnstr. 26, 731 (1955). 6 J. Durand, Rev. Sci. lnstr. 30, 840 (1959). 7 P. H. Cross, Rev. Sci. lnstr. 32, 1179 (1961). 8 J. L. Carter, E. V. Edwards, 1. Reingold, and D. L. Fresh, Rev. Sci. lnstr. 30, 946 (1959). 9 J. G. Grumberg, G. Antier, and P. E. Seiden, Rev. Sci. lnstr. 32, 979 (1961). 10 E. Heintzelman, J. G. Stewart, and C. G. Reed, Rev. Sci. lnstr. 33, 570 (1962). II Dry diamond powder may be obtained from the Industrial Diamond Powder Company, Pittsburgh, Pennsylvania. The epoxy used was CIBA Company's Araldite 502 with Araldite HN-951 hardener.\nVariable Frequency Power Source for Synchronous Motors\nP. E. V. SHANNON\nU. S. Naval Research Laboratory, Washington, D. C.\n(Received 31 July 1963; and in final form, 4 October 1963)\nI T is often desirable to change or vary the speed of a motor-driven device. A simple way to accomplish this for synchronous motors is by use of a variable-frequency power oscillator. This article describes such an oscillator capable of delivering 20 W at 115 V ac at any arbitrary frequency in the range from 25 cps to 12 kc. It will match correctly a variety of impedances, has good waveform when\n~687\nc ~t- '''~''WO'' , , I 02 I\n\" \"I \",.1\"'1 rF OSCILLATOR VOLTAGE FEED-BACK CONTROL ADJUSTMENT\nconnected to a motor load, has a high degree of frequency stability, has provision for intermittent on-off or remote operation, and has a finely adjustable frequency (i.e., speed) control.\nThe first stage (Fig. 1) is a \\Vien bridge oscillator employing a 6CG7 double triode. The frequency delivered by the bridge is given by the relation\nfrequency= 1/ (27rR tC),\nwhere C is the capacitance of a bridge ann, and R t is the total resistance in a bridge arm. In this case, R t is the sum of the fixed resistance Rl and the variable resistance Rv The purpose of the fixed resistance in the bridge arms is to compress the range of variation and thus provide finer frequency control. If desired, several ranges can be in cluded by providing the appropriate switching. Compo nents in this stage should have good temperature stability and be mechanically rigid. The voltage control is at the output of this stage and is coupled to a type 6072 double triode, one section of which is used as a voltage amplifier, and the other as a split-load phase inverter. The latter, in turn, drives a 5687 push-pull low impedance driver. Triode-connected 6550 beam power tubes are used in the final stage. These tubes are cathode biased, and on-off operation is accomplished in this stage by \"keying\" the cathode, thus avoiding the high voltage transients that would be developed in the output transformer if switching were done at the output. The output transformer is a United Transformer Company type CVM-1 multi-tap modulation transformer. This transformer was selected be cause of its variety of available impedances, wide fre quency response, and power-handling capabilities.\nThe power supply is a standard full-wave choke-input type. It is desirable to use choke input since keying the cathode of the power stage removes the major portion of the power-supply load that, with a capacitor-input supply, would cause the voltage to soar nearly to the peak trans former voltage and thus require the use of 700-V filter capacitors, which are bulky and expensive.\nGeneral construction is not critical, and the use of triodes provides a degree of stability not readily obtainable with pentodes. In line with good construction practice to prevent unwanted feedback, parts layout should follow the general geometrical arrangement of the schematic.\n6550\nII~ FIG. 1. Schematic of variable-frequency power oscillator. All capacitances are in.uF.\nThis article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitationnew.aip.org/termsconditions. Downloaded to IP:\n130.209.6.50 On: Thu, 18 Dec 2014 21:59:37", + "The only pre-use adjustment necessary is that of the oscillator feedback control. This control is adjusted for least distorted output waveform, or in the absence of distortion-analysis equipment, minimum amplitude out put at which stable oscillation still exists.\nThe over-all long-term stability, as measured with a Hewlett-Packard crystal-reference counter (model 521C), is better than 0.1%, which is greater than required for many applications.\nFairchild Semiconductor, Division oj Fairchild Camera and Instrument Corporation, Palo Alto, CaliJornia\n(Received 24 September 1963)\nIN the course of our work attempting to anodically oxidize single-crystal silicon slices, several technical problems have been encountered. One of the most import ant of these, a problem with which this note is concerned, was the development of an electrode assembly that would enable the immersion of the silicon anode into the anodizing electrolyte in such a way that the whole wafer would be immersed, but allowing an electrical contact to it with little or no current leakage. An additional desired char acteristic of the electrode assembly was its ability to be used in such a manner as to eliminate or minimize the preparatory steps of the silicon wafer anode.\nIt should be mentioned here that a number of methods of securing a silicon anode to the electrode assembly have been discussed in the literature'! These methods could be generally classified into three categories: (a) Thick pre formed oxide on the nonworking area; where the wafer is immersed into the electrolyte up to or a little over the\nedge of the preformed oxide so that this oxide acts as an insulator and presents electrolyte creepage; (b) \"stopping off\" compounds; where compounds such as Teflon and Apiezon waxes are applied onto the nonworking areas of the wafer to form insulators and creepage barriers; (c) Mechanical seals; where several rubber, Teflon, or other inert compounds are used to form mechanical pressure sealed insulators. The first two classes require, by their natures, a number of preparatory steps such as growing a layer of oxide or the application of the \"stopping-off\" compounds. These steps are t;me consuming and not al ways very successful. We believe that mechanical seals offer a simpler and more practical solution to the problem of securing the semiconductor anode to the electrode assembly since no special preparatory procedures of the anode are required, and the results obtained from such an assembly are at least comparable to the ones obtained with the other two classes of assemblies.\nThe electrode assembly, shown in Figs. 1 and 2, is made from two pieces of \"F\" shaped polypropylene cut so that they can be pivoted around a common point. The top end is spring loaded in order to keep the bottom arms closed. In each piece a vertical hole l2-in. in diam is drilled from the top of the clip to within t in. of the bottom edge. The top end of the hole is tapped to receive a screw. A horizon tal hole is then drilled in the bottom end to meet the vertical hole. A second larger horizontal hole is drilled around the first hole part way into each arm so that the conductor may be enclosed in a Silastic rubber sleeve. A spring is placed around a gold tip and placed in the horizon tal hole. The assembly is held in place by a tight fitting Silastic rubber sleeve. A l2-in. copper rod is placed down the vertical hole until contact is made on the side of the horizontally placed spring. This rod, which is held in place by the screw, forms a contact with the spring and gold tip and serves to hold the spring in place. Electrical contact is made to the top screws.\nThis article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitationnew.aip.org/termsconditions. Downloaded to IP:\n130.209.6.50 On: Thu, 18 Dec 2014 21:59:37" + ] + }, + { + "image_filename": "designv11_92_0003675_022092-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003675_022092-Figure8-1.png", + "caption": "Figure 8(a) shows the distribution of the lubricating oil film in the bearing. Red represents the lubricating oil, green the transition state, and blue the solid area. Although the applied external load is constant, each rolling element is subjected to different contact loads as the bearing operates. The rolling element with smaller contact load and that with larger contact load are selected to obtain the oil film thickness cloud map. As shown in Figures 8(b) and (c), on the rolling element with a smaller contact load, the minimum oil film thickness on the outer ring raceway is significantly larger than that of the inner ring. It is the same with the rolling element with larger contact load. Therefore, although the load on the inner and outer ring races is unevenly distributed, different rolling elements show the same trend of change in terms of the oil film thickness between the inner and outer races, that is, the oil film of outer ring contact race is thicker than that of the inner ring contact raceway. From the sectional view of the oil film thickness in Figure 8(d), it can be seen more clearly that the oil film thickness is distributed in a rectangular shape, and the rectangular area of the oil film thickness between the rolling element and the inner ring is smaller than that between the outer ring and the outer ring. .", + "texts": [ + " (2) At the inlet of the bearing lubricating oil, the thickness of the oil film drops sharply due to heavy bearing load; when the oil film pressure distribution is abruptly separated from the corresponding position of the Hertz curve, the oil film thickness begins to decrease, and the corresponding oil film becomes thinner. In the exit area, the thickness of the oil film increases sharply. Through the basic trend of thickness and pressure of oil film of bearing lubrication as shown in the figure, the simulation results are reasonable and correspond to the theory. Stimulate and analyze the thickness of contact oil film between the oil film and the inner and outer rings of the bearing under the working condition of the rotation speed of 8.714 rad/s. Results are shown in Figure 8. (a) Oil film thickness the whole image CISAT 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1168 (2019) 022092 IOP Publishing doi:10.1088/1742-6596/1168/2/022092 4.2 Oil Film Pressure Contact Change Analysis CISAT 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1168 (2019) 022092 IOP Publishing doi:10.1088/1742-6596/1168/2/022092 Stimulate and analyze oil film pressure of the oil film and the inner and outer rings of the bearing under the working condition of the rotation speed of 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002242_978-3-319-99262-4_29-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002242_978-3-319-99262-4_29-Figure5-1.png", + "caption": "Fig. 5. Seals housing detail", + "texts": [ + " This section describes the experimental setup for analyzing rotor seal systems at the Chair of Applied Mechanics at the Technical University of Munich, as in [21]. First, the seals test rig is presented, then the dynamic behavior is analyzed and compared to the simulation results. The experimental analysis is examined on the seals test rig (see Fig. 4). The main components are a flexible shaft and a mass disk (1) with two symmetricallyarranged liquid annular seals (2) in the middle (see details in Fig. 5). Eddy current sensors for measuring the displacement (6) and a piezo force platform (7) are arranged in the seals stator housing (8). The fluid is injected between the two seals with a maximum pressure of 100 bar. The rotor runs at over-critical speed above the first (sealless) natural frequency \u03c91 up to 12,000 rpm. An active magnetic bearing (3) is used as an exciter (2D shaker) at the shaft. The rotor shaft is supported by two ball bearings (4) and driven by a servo motor (5). The typical test procedure is the stationary rotor run-up (discrete rotational speeds) with or without AMB excitation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000265_omae2008-57626-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000265_omae2008-57626-Figure1-1.png", + "caption": "Fig 1. \u2013 3D figure for the proposed wave energy device composed by body (a) and (b) that include floaters (1), (2) and (3).", + "texts": [ + " Harren studied multiple rafts of varying lengths in shallow water as well as a three-raft system in shallow and deep water waves. From the study of the contouring raft conversion system two major problems were identified. First, the sea area required for the deployment of an array of such devices to supply a coastal town of 1000 homes from a 16 kW/m wave energy resource is too large (about 5 10 m2) [9]. The second major problem concerns the mooring of the raft. The proposed system aims to extract energy from sea waves by the relative pitch motion between bodies (a) and (b) (see figure 1). This relative motion is used to actuate the hydraulic rams that are inside floater (3). The installed power for the device is 1.2 MW. Since body (a) and (b) are geometrically different, the hydrodynamic characteristics of both bodies are also different. Thus the device is able to extract energy for a broad incident wave frequency band. In the next section a frequency domain model is developed which takes into consideration the heave, pitch and surge motions for the bodies, added mass and hydrodynamic damping as well as load damping for the power take-off mechanism. A time domain model is then devised which allows the non linear behaviour of the power take-off to be taken into consideration. Results are obtained for both models and average power available to the hydraulic motor is computed. The proposed system is presented in figure 1. As it is possible to see the device is made of two different bodies, (a) and (b) that can rotate around the longitudinal axis of the floating cylinder (3). There are three different floaters (1), (2) and (3). Body (a) and (b) have respectively a mass of 3.38 10 kg and 3.11 10 kg. The device has a total length of 70 m (35 m for each body). Floaters (1) and (2) have a draught of 12 m and a length of 22 m. Floater (3) has a draught of 4.5 m and a radius of 5m. Floaters (1) and (2) have 17 meters height", + " The volume displaced by these floaters at equilibrium is 2586.7 m3. The bridges (that link the central floater to the other two) have an equilibrium displaced volume of 245 m3 and 451m3 for body (a) and (b) respectively. As usual, it is assumed that the bodies have linear hydrodynamic behaviour. Thus diffraction and radiation forces and torques can be computed from linear hydrodynamic theory. For the frequency-domain model it is assumed that the power take-off equipment has a linear behaviour and can be simulated by a linear damper. Figure 1 shows the coordinate system and the two bodies (a) and (b) that can oscillate pitch, heave and surge. The pitch motion is defined as oscillatory rotation around the y-axis, i.e., the longitudinal axis for the floating cylinder (3). Heave and Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Term surge motions refer to z-axis and x-axis translations respectively. Due to symmetry, sway, yaw and roll motions do not influence the pitch motion of bodies (a) and (b). Thus for the considered coordinate system it is assumed that we have the following degrees of freedom, x for heave, for body (a) pitch motion, for body (b) pitch motion and d for surge", + " (15) The equation for the hydraulic motor flow rate must take into consideration the machine control strategy to be used. Following Falc\u00e3o in [1], we assume that \u2206 , (16) where is a control parameter. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/27/2016 Ter Solving equations (8-11) and (13-16), knowing that, , with , and , and also assuming that the exciting forces are known, it is possible to numerically simulate the device behaviour for irregular waves. The hydrodynamic coefficients for the wave energy device that is presented in figure 1 were computed with WAMIT\u00a9 software. Excitation forces for waves with 1m amplitude, hydrodynamic added mass and damping coefficients were computed. These coefficients were obtained for 131 angular frequencies/periods equally spaced from 0.05rad s-1/125.66s to 1.257rad s-1/5s. The water depth is 50m. For the frequencydomain analysis, only damping coefficients for the power takeoff equipment were needed. Constraints were imposed on the amplitude for the relative pitch motion between the two bodies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003917_978-3-030-12346-8_35-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003917_978-3-030-12346-8_35-Figure5-1.png", + "caption": "Fig. 5. Exit angle analysis by SolidWorks 2016", + "texts": [ + " Assuring a complete and correct mold filling is essential to avoid part defects. This can be achieved with a proper calculation of the feeding channels and risers, among others [9]. For the design of all the patterns, the SolidWorks 2016 software has been used. It can be especially highlighting the tool \u201cExit Angle Analysis\u201d, a graphical tool that allows to detect easily the pattern faces that needs an exit angle. These angles are useful to make easier the pattern extraction from the sand mold. Figure 5 shows the process with which the exit angle for the crankshaft is provided. Due to the large number of faces perpendicular to the partition line, the software marks in yellow the ones that need an exit angle. All patterns, except the piston, have been divided in 2 parts (cope and drag) and a keyway is designed to make easier their attachment, following the usual sand casting procedure. To build the optimal part shape, an approximation to the different analytical methods of calculation for the sand casting elements has been made" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002029_ilt-11-2016-0283-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002029_ilt-11-2016-0283-Figure2-1.png", + "caption": "Figure 2 A blowup view of the bump foil in Figure 1", + "texts": [ + "eywords Optimization, Optimal design, Bearings, OpenMP, Computational fluid dynamics, Fluid-film lubrication, MATALB, Parallel computing Paper type Research paper b = local best merit of the PSOmethod; B = bearing length; C = constant in truncation error term [Equation (3)]; c = radial clearance of foil bearing; c1, c2, c3 = parameters used in equation (7); e = eccentricity of foil bearing; g = global best merit of the PSOmethod; h = film thickness; h0 = characteristic film length of bearing; hmin = minimum film thickness; h = dimensionless film thickness; l = length of half bump foil; m = number of search steps (PSOmethod); n = number of grids or number of the particles (PSO method); N = number of iterations; p = film pressure; p = dimensionless film pressure; pa = average film pressure; pa = dimensionless average film pressure; r = radius of foil bearing; r1, r2 = random numbers [Equation (7)]; s = particle position in the PSOmethod; S = speedup factor [Equation (6)] or pitch of bump foil (Figure 2); tB = thickness of bump foil; tP = elapsed execution time in parallel computing (s); tS = elapsed execution time in sequential computing (s); U = bearing velocity; x, y = Cartesian coordinates; x; y = dimensionless coordinates; a = constant in equation (3); b = aspect ratio of bearing projected area; g = computing performance indicator [Equation (5)]; m = lubricant viscosity; v brg = rotational speed of foil bearing; v opt = optimal relaxation factor [Equation (3)]; Dp = residual of dimensionless film pressure; and Dx; Dy = dimensionless grid sizes in the x - and y-direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002994_fpmc2018-8876-Figure20-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002994_fpmc2018-8876-Figure20-1.png", + "caption": "FIGURE 20. COMPARISON OF SELECT DESIGNS", + "texts": [ + " However, two alternative designs were found using parallel coordinates charts that significantly reduce the value of OF3 while either improving or allowing for only a marginal increase in the other OFs. Both alternative designs were selected to give a good compromise of each OF value while maintaining or decreasing the size of the respective pump from literature. The proposed design alternatives are thought to be an improvement and are shown in Fig. 19. A comparison of the normalized OF values for the literature designs, their alternatives, and the selected design is shown in Fig. 20, while the input variables and OF values for the alternative pumps are shown in Tables 3-4 respectively. Comparison to the profiles found in literature demonstrates the value of an extensive multi-objective optimization of profile geometry as well as the necessity of a metric to quantify contact force, which was neglected in the work that determined the benchmark profiles and led to the high values of OF3 [26]. The relationships of each of the input variables on the OFs can be visualized using the scatter matrix shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003825_1754337119831107-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003825_1754337119831107-Figure8-1.png", + "caption": "Figure 8. Arrangement for measuring the pressure distribution around the 4.5-mm-diameter spherical lead pellet. (a) Spherical pellet (size exaggerated for clarity). (b) Fixed axis of rotation through pellet centre. (c) Stainless steel two-part tubular sting (0.8 mm and 2 mm o.d. tubes). (d) Flexible rubber tube to pressure transducer. (e) Sliding point for checking pellet rotation axis. (f) Frame for rotation of pellet about fixed axis b. (g) Stationary protractor for setting pellet rotation angle. (h) Flow nozzle. (i) 0.2 mm pressure tapping hole on spherical pellet surface.", + "texts": [ + "g. in the case of surface band \u2018d\u2019, pressure ps =pd = (p4 +p5)=2, where p4 and p5 were the pressures measured at pressure tappings 4 and 5 on either side of band \u2018d\u2019). A somewhat different experimental arrangement was used to obtain the pressure distribution around the 4.5mm spherical lead pellet (numbered 1 in Figure 2). This arrangement took advantage of the perfect symmetry of the spherical pellet and it involved only a single spherical pellet with a single pressure tapping hole. As shown in Figure 8, a 0.2mm diameter, pressure tapping hole was drilled on the surface of the spherical pellet. The hole was arranged so that it communicated via a stainless tubular sting and flexible rubber tubing with a pressure transducer. The pellet could be rotated, by means of a supporting frame, about a fixed transverse axis which passed through the centre of the spherical pellet. The tapping hole could be positioned at any azimuthal angle uo on a 180o semi-circular arc, which extended from the stagnation point, on the front face of the pellet when facing the flow, to a diametrically opposite point at the centre of the pellet rear face, as shown in more detail in Figure 9", + " This range is also within 4% and 8% of values obtained by Jourdan et al.43 in non-stationary flows using accelerating and decelerating spheres with Ma values of 0.47\u20130.61 relative to the flow. Overall, good agreement was found between the measured values of Cd for the nine spheres and those from published sources. Figure 9 shows the distribution of the pressure coefficient around the surface of the 4.5mm spherical lead pellet (No. 1 in Figure 2) that was measured as described in the previous section using the arrangement shown in Figure 8. The values of pressure coefficient shown in Figure 9 were calculated using the pressures measured at the pressure tapping hole, when the hole was located successively at various azimuthal angles, uo, between 0o and 180o (see Figures 8 and 9). Figure 9 shows that the minimum pressure coefficient occurred at azimuthal angle uo of just less than 70o and that flow separation occurred at an angle of just less than 80o, where the pressure coefficient recovery faltered due to the flow having separated from the pellet surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001431_6.2018-1286-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001431_6.2018-1286-Figure10-1.png", + "caption": "Fig. 10 Free Body Diagrams of Descending and Ascending Gliders", + "texts": [ + " Kirchoff modeled the added mass effect of a body in a fluid as additional mass in the kinetic energy equation [20]. Equation 1 is the kinetic energy due to the added mass effect. Mf , Jf , and D f Are the added mass matrix, added inertia matrix, and added cross term, respectively. v and \u2126 are the translational and rotational velocities. This analysis uses a body-fixed coordinate system aligned with the major and minor axes of the vehicle. \u03b8 is the glider pitch relative to horizontal, as shown in figure 10. The glide angle, \u03be, describes the angle that the path of the glider makes relative to horizontal. The angle of attack, \u03b1, is the angle the wing makes with the glide path. The angle of attack is positive for diving glides and negative for climbing glides [18]. Tf = 1 2 [ v \u2126 ] [ Mf DT f D f Jf ] [ v \u2126 ] (1) The free body diagrams in figure 10 show the forces and moments acting on a glider in a) dive and b) climb. Equation 2 is the force equilibrium equations in the horizontal and vertical directions of the global frame derived from figure 10.[ 0 m0g ] = [ cos(\u03be) sin(\u03be) \u2212 sin(\u03be) cos(\u03be) ] [ KD0 + KD\u03b1 + KD2\u03b1 2 KL0 + KL\u03b1 + KL\u03b1 ] V2 (2) Fig. 11 Camber/Reflex measurement The performance gains of this morphing wing can be quantified as energy saved during operation due to the reduction in movement of the trim mass. Typically, gliders move trim masses within their hulls to create the necessary pitch changes for climb-dive glides. The moment generated by a cambered wing will meet some of that need without expending electrical power. A possible drawback is the added drag of a cambered wing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003221_iecon.2018.8591476-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003221_iecon.2018.8591476-Figure1-1.png", + "caption": "Fig. 1. (a) Stator, rotor, and (b) cross-sectional view of the prototype single-phase DCFSM", + "texts": [ + " Subsequently, a simple search procedure uses the proposed model and the inductances calculated by FEA software to identify the optimum current profiles for different required torque with minimum copper loss. Finally, the proposed current profiles are first evaluated by FEA software, and then, it is validated experimentally through the prototype single-phase DCFSM. The simulation and experimental results both demonstrate that the torque ripple is greatly reduced when the motor is excited by the proposed current profiles. II. PROTOTYPE SINGLE-PHASE DCFSM Figure 1 (a) shows the stator and rotor of the prototype motor. As listed in Table I, the prototype motor is rated for 400 W and 0.64 N-m. The prototype motor used in this paper is an eightslot, four-pole single-phase DCFSM with a special two-layer rotor design along the axial direction. With such rotor design, the reluctance torque is enhanced at the positions where the motor cannot deliver sufficient initial torque. Therefore, the 978-1-5090-6684-1/18/$31.00 \u00a92018 IEEE 565 motor has a much higher starting torque about 50% of rated torque and can start from any positions with a simple starting procedure. The cross-sectional view of the prototype motor and definition of the rotor position \u03b8m are illustrated in Fig. 1(b). For a magnetically-linear single-phase DCFSM, its electromagnetic torque is determined by following equation mrelucArelucF mmfamf f ma a me TTT MiiM i M i T 22 22 (1) where mas m ma L d d M (2) mfs m mf L d d M (3) mm m mm L d d M (4) where ia and if denote the armature and field current, respectively. Las and Lfs denote the self-inductance of the armature and field winding, respectively. Lm denotes the mutual inductance shared by both windings. Note that the inductances are typically the function of the rotor position in a single-phase DCFSM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001642_j.ejc.2018.02.023-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001642_j.ejc.2018.02.023-Figure12-1.png", + "caption": "Fig. 12. N1 is foldable into an envelope.", + "texts": [ + " \u25a0 Let N be a Conway tile and let {v1, v2, v3, v4} be a 4-base of N. LetP = v1v2v3v4 be a parallelogram formed by a 4-base of N. Then it is easily verified, by using an equi-rotational transformation for Conway tile N, that N is foldable into an envelope if there exists P such that P is divided into two identical right triangles by its diagonal (Fig. 10(a)\u2013(c)). We call such a Conway tile N bi-right. Consider a Conway tile N1 in Fig. 11. Although N1 is not bi-right, but it can be folded into an envelope (Fig. 12). This fact can be explained as follows: 1. N1 is equi-rotational into a parallelogram stripS by hinging pieces at three points of its 4-base {v1, v2, v3, v4} (Fig. 13). 2. S can be folded into an envelope E whose corners are v1, v2, v3 and v4 (Fig. 14). 3. By Corollary 1, N1 is also foldable into the same envelope E. The paper [1] studies how many ways there are for a given strip to be folded into different rectangle dihedra. On the other hand, in this paper we focus on positioning 4-bases (4 corners of dihedra) of foldable strips into rectangle dihedra (Lemmas 1 and 2) and finding more kinds of figures which are foldable into rectangle dihedra (Theorems 3 and 4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003073_012066-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003073_012066-Figure1-1.png", + "caption": "Figure 1 Technical drawing of small tensile specimens (mm)", + "texts": [ + ", Dassault Systemes, Concord, MA, USA) and imported to Materialize (Leuven, Belgium) software for final orientation to run in ConseptLaser the M2 Cusing (ConceptLaser, Lichtenfels, Germany) 3D printer. The thickness of the samples are designed to be 0.5mm however, their printed parameters might be slightly different due to the accuracy of the printing machine. The length of the samples is adjusted with 22mm and 7 mm thickness on the machine grip part. Parameter \u201ca\u201d has been changed the range of 0.15mm to 4.2mm and 74 samples were printed in order to reach a fine result, Figure 1. The cross-sectional area was calculated by multiple with parameter \u201ca\u201d and \u201cb\u201d. All samples were printed perpendicular to the platform surface (z-direction). Cross-sectional area (S0) of the samples are up to 2.4 mm2. 46 samples were printed orientation of ZXY and 27 samples were built with YZX orientation, Table 3. Illustrates the number of printed samples with building orientation respectively. 5th International Conference Recent Trends in Structural Materials IOP Conf. Series: Materials Science and Engineering461 (2019) 012066 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001986_s0025654418010065-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001986_s0025654418010065-Figure3-1.png", + "caption": "Fig. 3.", + "texts": [ + "4) for some value of \u03d5 different from \u03c0/2, then for a setting angle \u2212\u03d5 there exists a solution \u03c0 \u2212 \u03b1\u0304. Therefore, only positive values of a setting angle will be considered in the following. c) A special situation occurs for \u03d5 = \u03c0/2. In this case, there exist a pair of isolated solutions \u03b1\u03041 and \u03b1\u03042 = \u03c0 \u2212 \u03b1\u03041 and a range [\u03b1\u0303, \u03c0 \u2212 \u03b1\u0303] of non-isolated solutions (Fig. 2). MECHANICS OF SOLIDS Vol. 53 No. 1 2018 3.1. Constructing the Dependence \u03b1\u0304(\u03d5) We now consider dependence of solution of (3.4) on different values of angle \u03d5. Fig. 3 shows a qualitative diagram of the graphical solution of this equation for an idealized plate with HAQ. Thick line denotes the graph of K(\u03b1), the fine lines represent the graphs of the function tan(\u03d5 \u2212 \u03b1) for certain values of \u03d5. Let \u03d5 be somewhat bigger than \u03c0/2. Obviously, both isolated solutions will be displaced somewhat, and the value of \u03b1 increases. A set of uninsulated solutions disappears, isolated branch \u03b13(\u03d5) separates from its left edge, root \u03b12(\u03d5) moves to the right, approaching the value of \u03c0 for \u03d5 \u2192 \u03c0. With increasing \u03d5 roots \u03b11(\u03d5) and \u03b13(\u03d5) approach each other and, at some value of \u03d5 = \u03d5\u2217 merge into one \u03b1\u2217, and then vanish, which is reflected in the right-hand side of Fig. 4 a. The left part of the picture is completed symmetrically in accordance with property b). This \u201cprocessing\u201d of Fig. 3 and the use of formulas (3.3) allow us to construct the dependences u\u0304(\u03d5) and v\u0304(\u03d5) on steady-state rotation regimes. These dependencies are qualitatively represented in Figs. 4 b and 4 c. Points A, B, B1, C, C1, which will be useful in constructing phase portraits in paragraph 5, are marked on them. MECHANICS OF SOLIDS Vol. 53 No. 1 2018 3.2. Stability of Steady Regimes Entering, as usual, deviations from the steady-state values of the variables x = u \u2212 u\u0304, y = v \u2212 v\u0304, z = \u03b8 \u2212 \u03b8\u0304 = \u2212(\u03b1 \u2212 \u03b1\u0304) we can write down the first-approximation equations by virtue of equations (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001674_978-3-319-76138-1_3-Figure3.6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001674_978-3-319-76138-1_3-Figure3.6-1.png", + "caption": "Fig. 3.6 Example of a 1R2T robot: (a) Pose with force-closure (b) pose without force-closure. Gravity forces are required to balance the platform", + "texts": [ + " Different definitions have been introduced to characterize a pose [134] and the most important concepts are reviewed below. A pose (r,R) of a cable robot is said to be in wrench-closure [134, 183] or controllable [474] if for each wrench wP \u2208 IRn there exists at least one distribution of cable forces f \u2208 IRm such that AT(r,R)f + wP = 0 with f > 0 . (3.14) 1Li [286] only deals with the non redundant case r = 0, i.e. six cables and six degrees-of-freedom. In other words, every wrench wP can be balanced with positive cable forces (see Fig. 3.6a). In contrast, the structure equation for a pose (r,R) may be such that one can only find solutions for special wrenches wP \u2208 IRn (Fig. 3.6b). This property is sometimes called force-closure when cable robots with purely translational motion (2T and 3T) are considered.2 This concept of wrench-closure or controllability is a rather theoretical concept disregarding any technical upper or lower limits on the cable force beside forces being positive. The notion is a purely geometrical problem without the need for introducing parameters for the desired wrenches on the platform and for the feasible forces in the cables. The definition of wrench-closure yields some principle limits of cable robots even when we accept infinitely large cable forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002435_978-3-319-99522-9_3-Figure3.3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002435_978-3-319-99522-9_3-Figure3.3-1.png", + "caption": "Fig. 3.3 Universal Hooke joint", + "texts": [ + " The upper rotor 1b of length B1B2 \u00bc l2 is inclined with respect to 1a of a constant angle a and is attached to frame B1xB1y B 1 z B 1 , rotating around the joint B1 with the angle uB 10, the angular velocity xB 10 \u00bc _uB 21 and the angular acceleration eB10 \u00bc \u20acuB 10. It has the mass mB 1 and the tensor of inertia J\u0302B1 , with respect to his frame. Finally, the same cross 2b of mass mB 2 \u00bc mA 2 and a new tensor of inertia J\u0302B2 is attached to the frame B2xB2y B 2 z B 2 and rotates relatively about zB2 with the angle uB 21, the angular velocity xB 21 \u00bc _uB 21 and the angular acceleration eB21 \u00bc \u20acuB 21 (Fig. 3.3). Pursuing along two independent ways OA1A2 and OB1B2 of this spatial mech- anism, we define the following transformation matrices a10 \u00bc au10; a21 \u00bc au21a T a h1; b10 \u00bc bu10bah2; b21 \u00bc bu21h1; \u00f03:37\u00de where we denote the matrices au10 \u00bc rot\u00f0z;uA 10\u00de; au21 \u00bc rot\u00f0z;uA 21\u00de bu10 \u00bc rot\u00f0z;uB 10\u00de; bu21 \u00bc rot\u00f0z;uB 21\u00de aa \u00bc rot\u00f0z; a\u00de; ba \u00bc rot\u00f0y; a\u00de h1 \u00bc rot y; p 2 ; h2 \u00bc rot z; p 2 : \u00f03:38\u00de The angles of rotation uA 21, u B 10 and uB 21 can be computed as function of uA 10, starting from following kinematic relations of constraint b21b10 \u00bc aBAa21a10; aBA \u00bc 1 0 0 0 0 1 0 1 0 2 4 3 5: \u00f03:39\u00de We choose the following three independent equations tan\u00f0uA 21 a\u00de \u00bc tan a cosuA 10 tanuB 10 \u00bc cos a tanuA 10; sinuB 21 \u00bc sin a sinuA 10: \u00f03:40\u00de The conditions of connectivity of the velocities are given by the derivative with respect to time of the above matrix geometrical conditions (3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003470_amcon.2018.8614963-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003470_amcon.2018.8614963-Figure7-1.png", + "caption": "Fig. 7 Snapshot of solid cutting simulation", + "texts": [ + " Figure 5 displays the execution dialogue of the developed interface where the NCI file is converted into the NC program. Moreover, the NCI file can also be transformed to the process operation sheet indicating tooling, raw material and machining parameters, as shown in Fig. 6. A solid cutting simulation software package, VERICUT\u00ae is used to confirm the generated NC data. Given the raw material size, the specifications of the cutting tool, NC data, the type of controller, and the kinematics chain of an NC machine tool, it can interactively simulate the material removal process of NC data. Figure 7 shows that a table/spindle-tilting type machine tool is constructed in the simulation environment and the finished workpiece is verified. The generated five-axis NC data are further verified by performing an experimental trial-cut. Figure 8 presents the final machined workpiece demonstrating the effectiveness and feasibility of the developed conversion interface. 122 ISBN: 978-1-5386-5609-9 This study presents a postprocessor algorithm which converts NCI tool path generated by Mastercam into multi-axis NC program" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003146_022045-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003146_022045-Figure2-1.png", + "caption": "Figure 2. Roughness contact model between friction pairs", + "texts": [ + " To simplify calculations, a single control unit is analyzed and their structure parameters as follow in Table 1. When considering the random roughness on the surface of the friction pairs, the roughness on the inner surface of the cylinder and the outer surface of the piston is set to be \u03c31 and \u03c32 respectively. The gap between friction pairs is filled with the Newtonian oil. If the moving speed of the piston is U and the minimum gap thickness of the gap is h0, considering their direct contact, the roughness contact model is established as shown in Fig.2. IMMAEE 2018 IOP Conf. Series: Materials Science and Engineering452 (2018) 022045 IOP Publishing doi:10.1088/1757-899X/452/2/022045 Thus, the gap between the friction pairs is the thickness of the oil film in the flow field. When the roughness is not considered, it can be seen from Fig.1 that the thickness of the oil film can be expressed as 10 00 , , yxh yxhh h p (1) Where: \u21261 represents a non-diamond micro-textured region, and \u21260 represents a rhombic micro- textured region, consisting of 4 straight lines, namely: ab L ya L xbL ab L ya L xbL ab L ya L xbL ab L ya L xbL yx yx yx yx ) 2 () 2 (: ) 2 () 2 (: ) 2 () 2 (: ) 2 () 2 (: 4 3 2 1 (2) When considering the roughness, it can be seen from Fig. 2 that the actual oil film thickness becomes: 21 hhT (3) Assuming that the fluid can be characterized as Newtonian and slip phenomenon is not considered, the modified equation is given as follows by introducing the above parameters [9]: )(6)()( 33 xx h U y ph yx ph x s cyx (4) Where x and y are the pressure flow factors along the x and y directions, c is the surface contact factor and s is the shear flow factor. Their values can be found in reference [10]. The term \u03c3 is the comprehensive roughness and can be expressed by 2 2 2 1 (5) IMMAEE 2018 IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001574_1.5018474-Figure17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001574_1.5018474-Figure17-1.png", + "caption": "FIG. 17. Typical force analytical graph for a model sphere.", + "texts": [ + " But note that the decrease of the downstream peak Fd due to shock attenuation by the upstream particles because of energy dissipation is possible. (3) Both the increase in sphere number and the decrease in distance between spheres promote wave interference between neighboring spheres. As long as the wave interference times are shorter than the peak times, the peak Fd and Cd are larger than those of a single sphere. This work was supported by the Natural Science Foundation of Zhejiang Province (Grant No. LY17E060006) and the National Natural Science Foundation of China (Grant No. 51006091). Figure 17 shows a typical force analytical graph for a model sphere. The sphere (denoted by a filled black circle) is imposed by a drag force, Fd, along the x direction, a tension, C, along the respective halves of an aluminum-alloy wire (represented by two blue lines), and a downward (i.e., opposite to the z direction) gravity, G. As a result, it obtains an axial acceleration, a (i.e., in the x direction). It should be reasonable to assume that the wire works in a range of elastic deformation so that the normal strain, e, is proportional to C, since each side of the wire is only weakly constrained by the tube walls and no any appreciable plastic deformation of the wire can be observed in the present experiments indeed. Thus, it follows that e \u00bc l1 l0 l0 \u00bc C EA ; (A1) where E, A, l0, and l1 represent the Young\u2019s modulus, crosssectional area, and original and present one-half lengths of the wire, respectively. According to Fig. 17, two geometric relations can easily be found that sin h1 \u00bc S=l1; (A2) l0=l1 \u00bc cos h1 cos h2; (A3) where S is the x-directional displacement of the sphere; h1 is one angle between each half of the wire and yz plane, while h2 denotes another angle between the projection line of the wire on the yz plane and x axis. Moreover, the equations of particle dynamics for the sphere are given as 2C cos h1 sin h2 \u00bc G; (A4) Fd \u00bc ma Cx; (A5) where Cx is the x component of C. Considering Eq. (A1), it can be expressed as Cx \u00bc 2C sin h1 \u00bc 2CS l0\u00f01\u00fe C=EA\u00de : (A6) Substitution of Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002075_978-1-4419-0068-5_6-Figure6.17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002075_978-1-4419-0068-5_6-Figure6.17-1.png", + "caption": "Fig. 6.17 Experimental setup for testing diffusion characteristics of TiO2 membranes [95]", + "texts": [ + " To open the closed ends, a dilute hydrofluoric acid/sulfuric acid solution was applied to the barrier layer side of the membrane to etch away the oxide and the membranes were then rinsed with ethyl alcohol. The acid rinse is repeated until the pores are completely opened as seen in Fig. 6.16b. Without disturbing their flatness, the samples were crystallized in anatase phase by annealing in an oxygen environment at 280 C for 1 h [99\u2013101]. Diffusive transport through TiO2 nanotubular membranes were studied using the apparatus shown in Fig. 6.17. The membrane was adhered with a cyanoacrylate adhesive to an aluminum frame, and then sealed between the two diffusion chambers, A and B. Chamber A was filled with 2 mL of 1 mg/mL solute solution (either glucose, phenol red, BSA 304 6 Use of TiO2 Nanotube Arrays for Biological Applications or IgG) and chamber B was filled with 2 mL of DI water (for glucose diffusion) or PBS (for phenol red, BSA and IgG diffusion). The assembled setup was rotated at 4 rpm throughout the experiment to eliminate any boundary layer effects" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003560_ichve.2018.8642038-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003560_ichve.2018.8642038-Figure7-1.png", + "caption": "Figure 7. Electric field distributions on and around the insulator (d=300 mm, d1=349 mm, d2=240 mm)", + "texts": [ + "66 kV/cm, the lower air gap will not get breakdown. However, keep the distance of d at 300 mm, keep the distance of d1 at 349 mm, then change the distance of d2 to 240 mm, the floating potential on the external object is 65.57 kV, the average voltage gradient of the lower air gap is 5.55 kV/cm, which is very close to 5.66 kV/cm, the lower air gap is in a state of critical breakdown. Thus it can be assumed that the critical breakdown distance of the lower air gap is 240 mm, the simulation results can be shown in Fig. 7. When there is no external object, the maximum electric field strength around the insulator is 14.43 kV/cm, in the case of with an external object, the maximum electric field strength of the insulator reduced to 13.52 kV/cm, meanwhile, the electric field near the external object is obviously distorted, as shown in Fig. 7, the maximum electric field strength has reached 25.18 kV/cm, which may resulted in the air gap breakdown of the insulator. In order to confirm the critical breakdown distance of the lower air gap, keep the distance of the upper air gap at a value of 349 mm, keep the distance of the lower air gap at a value of 240 mm, change the distance d, which is between the centre of the external object and the centre of the insulator, the corresponding external object length, floating potential of the external object and the average potential gradient of the lower air gap are shown in Table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000626_j.1934-6093.2005.tb00236.x-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000626_j.1934-6093.2005.tb00236.x-Figure2-1.png", + "caption": "Fig. 2. Geometric interpretation of Lemma 2.", + "texts": [ + " If r = 0, for example, switching will happen when the signs of both CA 1B and CB are the same, regardless of the magnitude of U . A simple geometrical interpretation is given below to illustrate the stability condition. First, observe that | | ,r r r r r . To see this, consider the follow- ing inequalities. For 0 r r , || || || ( ) || || ||WX rq WX rq r r q WX rq | | || || | | || || | | || ||r r q r q r r q | | || ||r q , and for 0r r , || || || ( ) || || ||WX rq WX rq r r q WX rq | | || || | | || || | | || ||r r q r q r r q | | || ||r q . Figure 2 depicts various sets (Eqs. (17)-(19)), showing the sliding surface and the sliding region for a 2nd order system. The illustration shows that the sliding surface within the invariant set is also the sliding region. For a more general case where the open loop filter could be unstable, a state constraint is proposed in this paper to ensure the sliding condition. From Eqs. (10) and (11), the energy function Ve can be made to decrease if the magnitude of CAX is bounded. Assume that the bound is possible and is equal to " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001529_s12008-018-0471-y-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001529_s12008-018-0471-y-Figure12-1.png", + "caption": "Fig. 12 The route of simulation", + "texts": [ + " The entire wear process is discretized into a series of quasi-static analysis. The total number of oscillation cycles is L and the increment in number of wear cycles per step is l (the wear step). The L is discretized into n wear increments and the increment of wear depth h at each contact node then calculated incrementally. The volume wear rate kD is got by experiment. P is the contact pressure between the inner ring and the liner. ds\u2032 is the single sliding distance between the contact surfaces. Figure 12 is a schematic of ds\u2032. The ds\u2032 is calculate through the Eq. (11). ds\u2032 = \u03b8 180 \u00b7 \u03c0 \u00b7 r (11) where \u03b8 is oscillation angle, r is radius. The sliding distance ds of wear step is calculated by ds = l \u00b7 ds\u2032, So the h j in the process j th wear increment is calculated by the Archard model. h j = P \u00b7 ds \u00b7 kD (12) In the process of j +1th wear increment, h j+1 = h j + h j . The total wear depth h is calculated by accumulating the increment of wear depth. (2) Calculation of the wear increment of the contact node It is generally assumed that the wear is carried out in the normal direction of the contact surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003059_icepds.2018.8571782-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003059_icepds.2018.8571782-Figure2-1.png", + "caption": "Fig. 2. Scheme for measuring of filler wire feed speed:", + "texts": [ + " The sensor registers the images obtained, when the surface is illuminated by the LED; then, after image processing by the DSP, the movement is detected on the recorded frames. Then, via the USB interface, a serial data is transmitted to the computer. It contains information on the linear increments (\u0394X and \u0394Y) of speed along two coordinate axes \u2014 X and Y. The maximum frame rate is 12.500 frames per second. The described principle is widely used in robotics, instrument engineering, human interface devices (HID) and devices for tracking object movements and coordinate measurements [13,14]. III. HARDWARE AND SOFTWARE SOLUTIONS Figure 2 shows a functional diagram for measuring wire feed speed. The main element of this system is wire feeder SSJ-18 1 \u2013 drive roller, 2 \u2013 pinch roller, 3 \u2013 wire, 4 \u2013 coil with wire, 5 \u2013 guide rails, 6 \u2013 flexible hose, 7 \u2013 outlet nozzle, 8 \u2013 electron beam, 9 \u2013 surfaced layer (weld bead), 10 \u2013 reducer, 11 \u2013 DC motor, 12 \u2013 linear power supply, 13 \u2013 module ADC / DAC, 14 \u2013 computer, 15 \u2013 optical sensor [15]. Drive roller of mechanism 1 is mechanically connected to reducing gear 10. Wire 3 is pressed by means of mechanism 2 with a pressure roller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002894_ihmsc.2018.10168-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002894_ihmsc.2018.10168-Figure1-1.png", + "caption": "Figure 1. Three DOF helicopter experimental platform", + "texts": [ + " By using the classical PID[3-4] control method in the theory of traditional optimal control or linear quadratic regulator (two Linear Quadratic Regulator, LQR) [5]tracking controller design method, has a certain effect, but the adjustment of PID parameters depend on working experience, and the method of LQR high dependence on the shortcomings of the accuracy of the system model, to further enhance the control the control effect in a certain extent. Re modeling of the helicopter model through this paper, establish multiple linear models of the controlled object, and introduces the model switching idea to effectively deal with the system nonlinearity and uncertainty, and then combined with the LQR optimal control[6] design method of multi model LQR controller was verified by the experiment results, the comparison results show that this method can improve quickly the accuracy and system response. Figure 1 is a three degree of freedom helicopter model system developed by Quanser[7], Canada. Its three degrees of freedom are height axis, pitch axis and rotation axis. The kinetic model can be described by the equations of motion on these three axes. The power of the system is generated by the screw motor, and its thrust is approximately proportional to the voltage applied to it. Kf is the thrust constant of the combination of the motor helical propellers. V is the added voltage, and the thrust generated by the motor is (1) A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003145_032005-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003145_032005-Figure1-1.png", + "caption": "Figure 1. Typical high-voltage cable laying models under complicated working conditions.", + "texts": [ + " There are 6 outdoor terminals, 6 GIS terminals and 192 intermediate joints in total. The project has 9 high drop sections, and the construction condition is extremely complex. Based on the above analysis, a typical integrated project model can be obtained from reference to some regions according to the survey data. On the basis of the analysis of the model, the principle of the continuous laying process and the related simplified calculation method will be determined, so as to guide the practical application of such projects. The specific model is shown in Figure 1 below: The main characteristics of the model are as follows: A. The range of values for L2 is generally between 20 and 30m, and the section of the model usually has a drop of more than 30m, which needs to meet the continuous laying condition of single channel. B. The range of values for L1 is generally between 20 and 200m, and the model generally has a long stay tube length and a narrow connected well. Considering that the difficulty of the project lies in stay tube section (L1) and high drop section (L2), the following article will focus on these two sections", + "8 ( cos( ) cos( ) 2 2 ( (1 sin( ))) / tan 2 ( (1 sin( ))) / tan ) 2 F W L R R h R h R CD (5) According to the above formulas (1) to (5), the comprehensive tensile force can be obtained: + + + +T F F F F F AB EF CDBC DE (6) In summary, the comprehensive tensile force of all the stay tube section laying in complex conditions can be calculated according to the formulas derived above. Specifically, the maximum traction force required for laying high-voltage cables in different lengths and different types of stay tubes can also be calculated, with reference to the laying trajectory model given in the actual laying scheme. Consistent with what has mentioned previously, a simplified diagram of the relevant model will be given at first. In order to obtain a simplified model of the arc-shaped vertical laying under high-drop section, the High Drop Section portion in Fig. 1 is simplified to the model in Fig. 4 below. According to the force analysis, the required traction force can be obtained for each segment: Force in arc DC segment: IMMAEE 2018 IOP Conf. Series: Materials Science and Engineering452 (2018) 032005 IOP Publishing doi:10.1088/1757-899X/452/3/032005 2 2 9.8 [2 sin (1 )( cos )] 1c D WR T e T e (7) Force in CB segment: 9.8 2 ( cos sin ) 4 4B c BCT T W L (8) Force in arc BA segment: 2 2 9.8 [(1 ) sin 2 ( cos )] 1A B W R T T e e (9) The maximum allowable traction of the cable will be further considered: mT A (10) In the above formula, is the allowable traction strength of the conductor(N/mm2) and A is the cable conductor cross-sectional area(mm2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002434_s1068798x18080087-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002434_s1068798x18080087-Figure4-1.png", + "caption": "Fig. 4. Radial ball bearing.", + "texts": [ + " The rated contact angle \u03b10 (rad) of the unloaded bearing, with the selected axial clearance, is as follows, according to [4, 5] (1) where \u03b40 is the working radial clearance in the bearing (the initial radial clearance with corrections taking account of the seating of the bearing races on the shaft and in the housing and also the operational heating); rm = rin.ch + rex.ch \u2212 Dw is the distance between the centers of the external and internal raceway channels; rex.ch and rin.ch are the radii of the external and internal raceway channels, respectively; and Dw is the roller diameter (Fig. 4). \u03b1 = \u2212 \u03b40 0 arccos [1 /(2 )],mr 3 The working contact angle \u03b1 of the bearing is greater than the rated contact angle \u03b10 on account of the elastic contact strain \u0394\u03b1el due to the axial force Fa (2)\u03b1 = \u03b1 + \u0394\u03b10 el. RUSSIAN The axial force may be found from the formula [6] (3) where C* is the bearing\u2019s force characteristic, N (4) \u03be is the camber of the bearing raceway (5) Cex and Cin are constants, mm4/3/N2/3; Z is the number of rollers; \u03beex and \u03bein are the cambers of the external and internal raceways" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001794_1.5034601-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001794_1.5034601-Figure1-1.png", + "caption": "FIGURE 1. Typical layout of a gantry crane", + "texts": [ + "NTRODUCTION Among loading equipment, various track-mounted gantry cranes are widely used (Figure 1). In this paper two methods are used to find an optimal control force ensuring the damping of vibrations of a load being moved with a gantry crane from one point to another in a given time interval. The first method is based on the Pontryagin maximum principle [1], uses the calculus of variations and is known as the most common classical method for solving similar problems. The second method is based on the generalized Gauss principle [2] and uses some modern approaches to solving the control problems by the methods of non-holonomic mechanics [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001710_042093-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001710_042093-Figure3-1.png", + "caption": "Figure 3. Interaction of the elastic wheel and road surface and shear of normal load \u0420z varied when travelling across the cross-section in the contact area (shaded figure): 1 - tyre cross-section compression line, 2 - extension line.", + "texts": [ + " Structural tyre properties (ply, cord material, protector wear degree) influence only proportionality factor Hs. The method for assessing wheel rolling resistance coefficient involves approximation of real shear of normal wheel loading for a specific cross-section by branches of a hysteresis loop with parameters of the elliptic power model of non-elastic tyre resistance and determination of the center of gravity of the curvilinear figure [4]. Let us consider the scheme of interaction of the elastic wheel which rolls without skidding and smooth non-deformed road surface (Fig. 3). Let us assume that the real wheel is flat. As a result, acting forces and torques have adjusted values; radial tyre deformation area is a contact area (for the flat wheel with lk). Changes in real shear (lines 1 and 2) of tyre cross-section loading relative to the ideal one (dashed lines) typical for an absolute elastic body are due to hysteresis losses in a deformed tyre. So, resultant 4 1234567890\u2018\u2019\u201c\u201d Rz of normal responses of the support road surface will be shifted relative to the line of normal loading Pz and move through the center of gravity of the curvilinear triangle (a shaded figure)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001622_access.2018.2814681-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001622_access.2018.2814681-Figure2-1.png", + "caption": "FIGURE 2. Seal clearance and geometry of dimples in a textured seal.", + "texts": [], + "surrounding_texts": [ + "INDEX TERMS Friction coefficient, JFO cavitation theory, mechanical seal, operation parameter, optimum dimple area density.\nI. INTRODUCTION Mechanical seal with dimples is defined as processed micron-sized dimples with a certain shape and distribution on seal faces, whose original model was set up by Etsion and Burstein [1] in 1996. It provides a new choice for people to search for high cost-effective mechanical seal for its better sealing performance, simpler structure, lower price, and easier to change the shape and size of dimples. Thus, many theoretical and experimental studies have been carried out to investigate the effect of structure, distribution and geometric parameters of dimples on its performance in the last few years [1]\u2013[9]. See recent reviews in Ahmed et al. [10] and Gropper et al. [11].\nDimple depth, diameter, and area density are three basic parameters of evenly distributed dimple patterns. An objective study by orthogonal method indicated that area density is more significant for the tribological performance than the depth or diameter of dimples [12]. Many researchers have contributed to find the optimum dimple area density that minimizes the friction coefficient or maximizes the load-carrying\ncapacity of sealing surfaces through experimental study. Yu et al. [13] conducted a series of tests to investigate the effect of dimple area density on friction torque and face temperature rise of mechanical seal under various rotation speeds. They observed that both of the friction torque and face temperature rise are minimal when the dimple area density is 50%. The SiC disks with various dimple parameters were tested against a SiC ring by Wang et al. [14] under water lubrication. Their results show that there is an optimum dimple area density (5%) with sufficient water supply, where the critical load can be improved at least twice over that of un-textured surface. However, whenwater is not supplied sufficiently, dimples with higher area density (5\u201315%) are preferred. Kovalchenko et al. [15] experimentally investigated the effect of laser surface texturing on lubrication regime by measuring friction and electrical-contact resistance in a pinon-disk sliding conformal contact. They found that the pattern of dimples with lower area density (12%) is beneficial for expanding the range of the hydrodynamic lubrication regime. In the case of metallic seals, Qiu and Khonsari [16] evaluated\n14030 2169-3536 2018 IEEE. Translations and content mining are permitted for academic research only.\nPersonal use is also permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.\nVOLUME 6, 2018", + "the effect of dimple area density on friction coefficient by using textured steel rings. The experimental investigation shows that the friction coefficient decreases as the area density of dimple increases from 40% to 58%.\nOn the other hand, lots of theoretical studies which are thought to be the preferred to trial-and-error experimental works were carried out to obtain the optimum dimple area density. As early as in 1999, Etsion et al. [2] developed a model for mechanical seals with regular dimple patterns. Their simulation results suggest that the preferable percentage of area density is 20%, and above this value the rate of performance improvement becomes small. After that, a series models based on Reynolds equation have been developed by the same group. Usually, these analytical models suggest that the area ratio of 20\u201340% would be preferable since the total hydrodynamic pressure is maximized within this range. Du and Peng [17] applied the finite element method to solve Reynolds equation and investigated the effect of dimple area density on sealing performance. They observed that the peak value of opening force or film thickness can be obtained when dimple area density is selected between 40% and 60%. Kligerman et al. [18] found that the dimple area density should be largest in order to get higher friction reduction through several numerical simulations. Wang and Zhu [19] developed a virtual texturing technique for mixed lubrication regime in 2005. They found that in the case of low-speed/high-load, the lubrication performance is very sensitive to the variation of dimple area density. For the high-speed/low-load case, there seems to be an optimal range of dimple area density from 3% to 12%, where a dimple area density of 5% looks to be the best choice. A brief review on the studies related to the area density issue of dimple patterns can be seen in Wang et al. [20], where the design principles of dimple area density for mixed lubrication regime were also discussed.\nAs can be seen, different values of optimum dimple area density were achieved for the above existed studies. This originates from different operation conditions used in the studies and it indicates the great impact of working parameters on optimum dimple area density. On the basis of previous studies, a further in-depth research and exploration on the influence of operation parameters need to be conducted for the\nfollowing reasons: Firstly, Reynolds cavitation boundary condition was generally adopted when solving Reynolds equation in previous theoretical studies. However, it has been proved that this boundary condition is non-massconservative, and will give incorrect results [21]\u2013[23]. Moreover, the interplay between operation parameters was ignored for most of the previous studies. That is, keeping the other ones as constants when investigating the influence of one of the operation parameters. This may lead to different results if the constants change to other values.\nIn view of this idea, in the present paper, the influences of operation parameters including sliding velocity, outlet pressure, and seal clearance on optimum dimple area density for mechanical seal, are studied by solving Reynolds equation, where the JFO cavitation theory is implemented to take the cavitation into account. The interaction effects between the operation parameters are also considered.\nII. ANALYTICAL MODEL The mechanical seal considered in the present paper is represented by two non-contacting rings rotating relative to each other. The regular network of dimples is fabricated on the seal face of stationary ring, the geometrical model of which is shown in Figs. 1 and 2. The relationship between the inner and outer radii, ri and ro in Fig. 1, of the seal rings under consideration satisfies that (ro-ri)/ri 1. This allows one to neglect curvature effects and consequently, a circular sector containing one dimple column in the radial direction\nVOLUME 6, 2018 14031", + "is assumed to be rectangular with length l and width w, subjected in the lateral x direction to a relative sliding velocity u, corresponding to the tangential velocity at the mean radius of the seal. The dimples with number n are distributed uniformly over the column and the seal clearance is denoted by h0. Each dimple is located in the center of an imaginary cell with length l/n and widthw and modeled by an axisymmetric segment with a base radius rp and depth hp. The dimple area density Sp can be defined as\nSp = n\u03c0r2p lw\n(1)\nThe two-dimensional, steady-state form of the Reynolds equation for an incompressible Newtonian fluid in a laminar flow is expressed by\n\u2202\n\u2202x ( h3 \u2202p \u2202x ) + \u2202 \u2202z ( h3 \u2202p \u2202z ) = 6\u00b5u \u2202h \u2202x\n(2)\nwhere \u00b5 is dynamic viscosity, h and p are the local film thickness and pressure at a specific point of the computational domain, respectively.\nFor the dimple with elliptic section studied in this work, h can be given by\nh = h0 outside the dimple h0 + hp rp \u221a r2p \u2212 ( x \u2212 l 2n )2 \u2212 ( y\u2212 w 2 )2 inside the dimple\n(3)\nAs shown in Fig. 1, the following boundary conditions are applied\np(x, z = ri) = pin, p(x, z = ro) = pout (4a)\np(x = \u2212w/2, z) = p(x = w/2, z) (4b)\nwhere pin and pout are inlet and outlet pressures of the mechanical seal, respectively.\nThe above boundary conditions should be complemented by the conditions at the boundaries of possible cavitation regions associated with the dimples. In the present work, the Floberg\u2013Jakobsson\u2013Olsson (JFO) cavitation theory is implemented using a mass-conservative algorithm to accurately predict the behavior of cavitation. According to the JFO cavitation theory, the magnitude of the pressure in the entire cavitation region remains constant at the cavitation pressure pcav, which is \u2018\u2018predetermined\u2019\u2019. The film rupture starts at the location where the pressure derivative with respect to the normal direction is zero. The film reformation boundary is described by the following formula:\nh2\n12\u00b5 \u2202p \u2202n = Vn 2\n( 1\u2212 \u03c1\n\u03c1c\n) (5)\nwhere Vn is the fluid velocity in normal direction, \u03c1 and \u03c1c are the local and cavitation density of the fluid.\nBased on (2), to implement the JFO thoery, the universal form of Reynolds equation developed by\nElord and Adams [24] is adapted in the present paper, which can be written as [25]\n\u2202\n\u2202x ( h3 \u2202p \u2202x ) + \u2202 \u2202z ( h3 \u2202p \u2202z ) = 6\u00b5u \u2202\u03b8h \u2202x\n(6a)\np \u2265 pcav \u03b8 = 1 (6b)\np = pcav \u03b8 < 1 (6c)\nwhere \u03b8 (= \u03c1/\u03c1c) is the film content. By specifying the film thickness distribution h(x, z) and the relevant boundary conditions, the Reynolds Equation (6), is discretized by the finite difference method and then solved by the successive over-relaxation Gauss-Seidel iterative method for the pressure distribution p(x, z) in the film. Using the pressure results, the load-carrying capacity for the dimple column is calculated from the expression\nW = \u222b l\n0\n\u222b w/2\n\u2212w/2 p(x, z)dxdz (7)\nThe friciton coefficient is computed by consdering the total friciton force over the dimple column, which is given by\nF = \u222b l\n0\n\u222b w/2\n\u2212w/2 \u03c4 (x, z)dxdz (8)\nwhere the expression for the shear stress \u03c4 (x, z) of the lower surface is\n\u03c4 (x, z) = 0 p \u2264 pcav \u00b5u\nh (x, z) \u2212 h (x, z) 2 \u2202p (x, z) \u2202x p > pcav (9)\nThe resulting expression for the coefficient of friction is\nf = F/W (10)\nThe optimum dimpe area density in this article is defined as dimpe area density that minimizes friction coefficient.\nThe correctness and accuracy of the solution for (6) was verified with a problem taken from Qiu and Khonsari [26], where one dimple within a unit cell is considered. It was found that the pressure distribustions and the corresponding values of load-carrying capacity are nearly identical in the two studies [27].\nIn the present study, according to (1), the changes of dimple area density are obtained by changing dimple number n along the column while keeping dimple radius, column length, and width as constants. The influneces of three operation parameters, including sliding velocity, outlet pressure, and seal clearance are taken into account. To consider the interaction effects between the operation parameters, a total of 2500 simulation calculations are carried out to obtain the optimum dimple numbers (defined as n\u2217), corresponding to optimum area densities under various operation conditions. The values of the related parameters are listed in Table 1.\nThe choice of cavitation pressure pcav can strongly influence the simulation results from Elrod algorithm, which implements the JFO theory. According to the study of Shen and Khonsari [28], the simulation results using the\n14032 VOLUME 6, 2018" + ] + }, + { + "image_filename": "designv11_92_0001719_cyberi.2018.8337561-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001719_cyberi.2018.8337561-Figure2-1.png", + "caption": "Fig. 2. Drawing of the proposed model, weight (load) is on the right hand side of the lever and moving counterweight is on the right hand side", + "texts": [ + "00 \u00a92018 European Union Let us illustrate a simple situation when we know the weight m in distance from the axis of rotation x1 and the counterweight M. The distance x2 of the counterweight corresponding to a zero torque is then calculated from i = 0 (1) where i denotes torque corresponding to ith weight; in our case i=1,2 and 1, 2 correspond to m and M respectively. Substituting the corresponding torques to (1) we have M* x2 \u2013 m* x1 = 0 (2) x2 = (m* x1)/M (3) The distance x2 of counterweight M from the point of suspension of the arm is thus easily calculated. The same principle is used in our Balancing Arm Model (BAM), Fig.2, when equilibrium is achieved by changing the position of the cart (counterweight). Keeping equilibrium state is considered as the basic control aim. Equilibrium state is achieved by a suitable cart distance x2 from the rotation axis. Controller should provide the control action getting the cart into a position x2 given by (3). The distances determined by the calculation and positioning by the controller may differ slightly. It is caused by friction of the bearings in the axis of rotation. Larger friction causes larger difference in the real position of the balancing cart corresponding to equilibrium state of the real BAM and x2 obtained from (3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003907_978-3-030-11434-3_26-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003907_978-3-030-11434-3_26-Figure1-1.png", + "caption": "Fig. 1. Hand tremor patient model", + "texts": [ + " Patients with this disability will see their hands oscillating in the range of 4\u201312 Hz [1]. It is very difficult for the people suffering from hand tremor to perform daily tasks such as eating and drinking. Medical treatments and brain surgery can be applied to tremor patients. However, for people not responding well to medical treatments, mechanical vibration absorbers can be a good alternative [2\u20134]. In order to study the effects of vibration absorbers on hand tremor patients, a model of such a patient has been designed and constructed (see Fig. 1). A model, although incomplete, is very useful in the studying of a system and in predicting the effects of different components. To narrow the scope of the project, we designed the model to simulate a patient with hand tremor sitting at a dining table trying to eat with a spoon. ADC fan with rotating unbalance is mounted on the model to create the tremor. The benefit of using such a model to study a vibration problem is that very repeatable results can be expected. Also, since the model is made out of wood, it is very lightweight and portable", + " The absorber consists of a mass element, M2 mounted at the end of a stiffness element, K2, and a clamp to mount the absorber onto the system for vibration control. To understand why the excitation frequency needs to be close to the system natural frequency to work, it is useful to consider the phase angle information between the input and the absorber provided in Fig. 4. When the absorber mass is \u2212180\u00b0 out of phase with the input, cancelation will occur. Figure 5 shows the hand tremor displacement versus time graph for the patient model shown in Fig. 1 with the DC fan turned on. The speed of the fan was adjusted to 11 Hz to excite one of the vibration modes of the system. The amplitude of oscillation was recorded by an accelerometer and converted into displacement. As shown below, the displacement amplitude at steady state was recorded as 2.8 mm (or 5.9 mm peak to peak). It is apparent that a hand tremor patient with 5.9 mm peak to peak vibration amplitude at 11 Hz would not be able to eat with a spoon. A vibration absorber with natural frequency at 11 Hz was designed with a mass of M2 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001087_iros.2009.5354509-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001087_iros.2009.5354509-Figure11-1.png", + "caption": "Fig. 11. Experimental setup", + "texts": [ + " The stress at the fragile tissue did not largely exceed the limit stress in any cases of organ models (a)-(e). This result suggests that our control method successfully prevents stress overload. The proposed parameter setting method is effective in achieving robustness relative to the variation in the stiffness parameter of the organ. In this section, we evaluate whether our control method is effective to realize sufficient performance to the real hog liver, the stiffness parameter of which is unknown. A. Experimental manipulator Figure 11 shows the manipulator used in the experiment. The experimental manipulator has four degrees of freedom achieving planner movement. The manipulator used for this experiment consisted of two parts: namely positioning and pushing parts. A positioning part has three serial joints to help position the pushing part with three rotation degree of freedom. The pushing part only realizes the translation movement to push the organ. A six-axis force/torque sensor (NANO 1.2/1, BL AUTOTEC) is attached to the root of the pushing part and the surgical tool is attached to the six-axis force/torque sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.18-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.18-1.png", + "caption": "Fig. 18.18 Self-aligning bearing", + "texts": [ + " By contrast, the heat partition ratio \u03bb is very dependent on the kinematics, since the transient interaction is symmetrical with respect to the two bodies and hence [for example] with similar materials we obtain \u03bb=1 for all sliding speeds. Lee et al. (2016) also give the expression Q = \u03c1cpa 3(T\u221e 2 \u2212 T\u221e 1 ) ( \u03c0 \u221a 2 + 1.8 Pe Pe ) (18.129) for the total heat exchange between the sliding bodies through a single asperity interaction with maximum contact radius a. This result can be used, in combination with an asperitymodel theory of rough surface contact, to predict the effective thermal contact resistance between two sliding bodies (Liu and Barber 2014). 1. Figure18.18 shows the cross-section of a spherical roller bearing. The inner surface of the outer race has a spherical profile so that the whole inner assembly\u2014inner race + rollers\u2014can rotate about the bearing centre in the plane of the paper, permitting the bearing to act as a \u2018pin joint\u2019. Spherical bearings are therefore tolerant of misalignment and are used in situations where manufacturing tolerances or shaft and housing deflections make it difficult to maintain exact alignment. The spherical inner surface demands that the rollers be barrel-shaped as shown in the figure and there is therefore always slip at some parts of the contact area" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000315_13506501jet616-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000315_13506501jet616-Figure3-1.png", + "caption": "Fig. 3 Configuration of sensors distribution and pad number in the experimental bearing", + "texts": [ + " The lubrication system of the test rig supplies oil to each bearing with an independent oil supply system. Eddy current sensors with an output of 7.87 v/l mm are installed on each bearing housing. In the process of test, the sampling frequency is 32 times the rotating speed. To obtain the circumferential pressure distribution, six pressure sensors are installed on one pad and totally 24 sensors are used. The detailed process is presented in Fig. 2. The detailed layout of pressure sensors is shown in Fig. 3. The oil pressure distribution of bearing mid-plane is transferred by the channel shown in Fig. 4. As shown in Fig. 3, the AE sensor is mounted on the housing and two positions are used, AE sensor 1 is on the top of the housing and AE sensor 2 is on the lateral face near the lower pads. The AE sensors (piezoelectric-type, produced by the PAC company, WD series) with an operating frequency range of 100\u2013 1000 kHz are employed. The preamplifier ranged from 40 to 60 dB (bandwidth between 20 kHz and 1.2 MHz) to eliminate the low and high components of the signal. The signal output from the preamplifier is Proc. IMechE Vol", + " The steady-state analysis is achieved as follows: a certain rotating speed is kept by employing a 55 kW variablefrequency motor and then maintaining this speed for about 5 min. In this case, the results of \u2018speed raising\u2019 and \u2018speed reducing\u2019 are nearly the same. Therefore only the \u2018speed rising\u2019 results are presented. In the present study, a number of tests were carried out to verify the pressure at different rotating speeds and to investigate the impact between the pad and the rotor accurately. Most of the measurements were repeated several times and good reproducibility was found. The pad number distributions are presented in Fig. 3, where the pads with nos 1 and 4 are the upper pads and the pads with nos 2 and 3 are the lower pads. The rotating speed is clockwise. The pressure distributions of different pads under different rotating speeds are shown in Fig. 6. It depicts the pressure of test points on four pads (nos 1, 2, 3, and 4). In Fig. 6(a) the pressure distribution on pad no. 1 is shown. It indicates that on the test point no. 16 (corresponding to 45.6\u25e6 from the lead edge) the convergence wedge is formed for the reason that the pressure increases", + " Engineering Tribology at Monash University on June 17, 2015pij.sagepub.comDownloaded from turbine machinery monitor. Therefore a more efficient way to identify the impact phenomenon is needed. The AE signals with different rotating speeds obtained in this experiment are presented in Fig. 9. The peak amplitudes with different period time mean the impact between the pad and rotor, which is the direct cause for the failure of the tilting pad bearings. The signals are from the test position of A1, as shown in Fig. 3. It is observed that the signals of rotating speed which equals 1000 r/min (in Fig. 9(a)) are distinguished from the other three rotating speeds. That is because the amplitude of the signals from the impact is small and the noise can cover it. In Figs 9(b) to (d), it indicates that the signals of different rotating speeds have different period times and different maximum peak-to-peak amplitudes. First, the period time with different rotating speeds are compared. The period time of 0.08, 0.048, and 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003681_978-3-030-04975-1_89-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003681_978-3-030-04975-1_89-Figure7-1.png", + "caption": "Fig. 7. Stress and strain distribution in the modified shapes pump bodies a; b) external gear pump c; d) internal gear pump e; f) gerotor pump", + "texts": [ + " In that way, the area affected by high pressure was reduced, and, consequently, the forces deforming the body in axial direction A were reduced. A similar modification of outlet O was made in the gerotor pump (see Fig. 2g). The modified shape bodies of the three considered pump types were subject to the FEM strength analysis. The loads and restraints distribution, as well as the numerical models, were developed according to the rules similar to those for the basic shape bodies. Also, the research programme adopted for the modified shape bodies was the same as the one for the basic shape bodies. Figure 7 shows the results of the stress and strain analysis of the modified shape pump bodies loaded with torque T = 16 Nm, Q = 35 Nm and pressure P = 10 MPa. The analysis of the stress distribution shows that the greatest reduced stress r, as in the case of the basic shape, occurs on the outlet side of the bodies, in the area of outlet O. Treating stress r as the criterial stress, its interrelation to pressure p loading the pump, namely r = f(p) for three modified kinds of bodies was studied. The interrelation is shown in Fig", + " 6a, b, c in a form of the dashed line. Figure 6a, b, c show that the stress increases along with an increase in pressure p according to the direct proportional dependence for all the modified bodies. When assessing the state of stress, it is stated that for all bodies with modified shapes, when their working pressure is p = 0\u201320 MPa, the stresses do not exceed the permissible values. At the same time, the stresses were lower than the stresses in the basic body. Analyzing the state of deformation, shown in Fig. 7, it is stated that modified bodies, similarly to basic bodies, are deformed in the axial (longitudinal) A and radial (transverse) R directions. In Fig. 7a, c, e show that still the greatest axial deformation dA occur on contact surfaces C between middle and front body, as well as between middle and back body. In turn, Fig. 7b, d, f shows that the largest radial deformations dR occur in the middle body near outlets O. Figure 6 shows that axial dA and radial dR deformations also increase in direct proportion to the pressure p that loads the pump bodies. When assessing deformations, it is noted that in all bodies with modified shapes, axial and radial deformations are smaller than in bodies with basic shapes, in the range of the same loads p = 0\u201320 MPa. It was particularly advantageous to reduce radial deformations that fell below the permissible values, which can be written: dRE; dRI; dRG\\dRallow \u00bc 0; 040 mm Axial deformations dAE, dAI, dAG were also reduced, but to a lesser extent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003681_978-3-030-04975-1_89-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003681_978-3-030-04975-1_89-Figure1-1.png", + "caption": "Fig. 1. Classification f gear pumps by the type of gearing and body shape [38]", + "texts": [ + " The cooperation of these teams with the body will ensure the process of converting mechanical energy into a hydraulic one. The course and efficiency of this process depends on the shape and dimensions of the body. Therefore, the body should meet the strength conditions, i.e. the stresses in the body and its deformations should be less than the allowable stresses and deformations, i.e. r < rallow, d < dallo. From the technological point of view, the body should have a simple shape, and its implementation should be possible by available and cheap production methods. Figure 1 presents typical shapes of the three basic types of gear pumps: \u2022 external gear pumps \u2013 no. 1, 4, 7 \u2022 internal gear pumps \u2013 no. 2, 5, 8 \u2022 gerotor pumps \u2013 no. 3, 6, 9 An analysis of the examples presented of the pumps allows for the distinguishing of three typical body shapes: \u2022 a cylinder-shaped body \u2013 no. 1, 2, 3 \u2022 a rectangular base prism-shaped body with optional curves - no. 4 or a square base prism-shaped body \u2013 no. 5, 6 \u2022 irregular- or fancy-shaped body \u2013 no. 7, 8, 9 Cylindrical or prism body shapes meet the specified design and technological criteria" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.36-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.36-1.png", + "caption": "FIGURE 16.36 Illustration of the limiting dome height measurement. (a) Schematic of the limiting dome height test. (b) Plot of dome height at failure versus blank width. (c) Schematic of specimen after testing. (From Bayshore, J. K., Williamson, M. S., Adonyi, Y., and Milian, J. L., 1995, Welding Journal, Vol. 74, pp. 345s\u2013352s. By permission of American Welding Society).", + "texts": [ + " The test measures the dome height at the onset of necking in the material. It tests the material in or near plane strain, and the outcome is affected by work hardening and ductility of the workpiece, as well as the frictional conditions. It does not distinguish between the effects of the individual variables. Furthermore, it cannot be used to specify desired material characteristics. The rectangular specimen has two sides solidly clamped to a binder, while the unconstrained edges allow control of the strain state along the centerline (Fig. 16.36a). Specimens of different widths are used, to determine the minimum dome height at fracture, referred to as the limiting dome height, (LDH0) (Fig. 16.36b and c). Fracture occurs in or near plane strain. The LDH test has been found to correlate well with the Rockwell hardness. When the formability of a laser-welded blank and a single blank are compared using, say, the dome height test, it is normally found that the maximum dome height achieved for the laser-welded blank of materials of either the same strength or thickness is less than that of the single blank, and the height further decreases with an increase in either the strength or thickness ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003059_icepds.2018.8571782-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003059_icepds.2018.8571782-Figure1-1.png", + "caption": "Fig. 1. Electron beam freeform fabrication process scheme: 1 \u2013 cathode; 2 \u2013 heater; 3 \u2013 anode; 4 \u2013 control electrode; 5 \u2013 electron beam; 6 \u2013 magnetic lens; 7 \u2013 deflecting coils; 8 \u2013 wire; 9 \u2013 feed mechanism; 10 \u2013 area of beam\u2019s impact on metal; 11 \u2013 substrate; 12 \u2013 weld bead", + "texts": [ + " With using of this technology it is possible to manufacture metal parts of various shapes with the advantages of vacuum metallurgy. It is already clear that this technology will be used for small-lot and small-scale production of various parts in science-intensive industries \u2013 for example, in the aerospace industry. It is also suitable for the production of semi-finished products for further mechanical treatment. However, the control of this process is a very complex technical task. There is some technical developments in this field, such as systems with feedback [3- 5]. The process is based on electron-beam surfacing (Figure 1), in which electron beam moves rapidly along a circular or other closed trajectory and melts the wire that is fed continuously. An electron gun, consisting of cathode 1 made of lanthanum hexaboride, heated to working temperature of 1600\u20131800 K by electron bombardment from tungsten wire heater 2, anode 3 and control electrode 4, forms electron beam 5. The beam is focused by magnetic lens 6 and deflected along predetermined path by the coils 7. The wire 8 is fed into the area of beam\u2019s movement by means of feed mechanism 9. Electron beam interacts with This work was carried out in the National Research University \u00abMoscow Power Engineering Institute\u00bb; it was supported by grant from the Russian Science Foundation (project 17-79-20015). the wire (pos. 10 on the Figure 1) and also with the substrate 11 and weld bead 12. In the result, the built-up metal is forming by deposition of many weld beads building from remelted wire. The advantages of electron beam surfacing, in comparison with traditional methods, include: reducing the cost of materials, high speed of the process, simplicity of developing technical documentation. However, at present, there is no mass production of electron beam equipment for freeform fabrication. This is due to the high cost of existing equipment and the lack of unified technical solutions for control systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003990_j.enganabound.2019.04.008-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003990_j.enganabound.2019.04.008-Figure11-1.png", + "caption": "Fig. 11. Assembly drawing of the bottom roll.", + "texts": [ + " 10 shows the calculation program flowchart of the Bearing BEM nder the EHL condition. .3. Numerical example In this paper, the load and pressure distributions of four-row cylinrical roller bearing will be analyzed by the bearing BEM under the EHL ondition, which will prove the validity and effectiveness of the method. .3.1. Computation models A four-row cylindrical roller bearing is taken as the research object, hich is installed on the operating side of the lower backup roll of a wo-roll experimental mill. The assembly drawing of the bottom roll is hown in Fig. 11 . The four-row cylindrical roller bearings are used to ear radial loads, and the single row tapered bearing and the two-row hrust spherical roller bearing are used to bear axial loads. The operating side of the lower backup roll system is taken as the alculation model. The elasticity Modulus of the mill roll and shaft block oth are 210 GPa, and the Poisson ratios both are 0.3. The structure and eometry of the calculation model have been simplified as follows: (1) The rollers are simplified to plate units and fixed on the outer race" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002493_s11277-018-5949-1-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002493_s11277-018-5949-1-Figure2-1.png", + "caption": "Fig. 2 To create the regional coordinate system indicated by landmark , the robot approaches to a (c) location that is a fixed distance from and in front of the landmark", + "texts": [ + " Dead reckoning resulted in straight-line motion for localization computation. Because the robots location was estimated by the surrounding observed landmarks and their spatial relations, a more accurate estimation of the location was produced for iterative feedback, and the accumulated error was thus minimized. The algorithm of semantic representation is as follows: 1. When the robot begins, the egocentric-semantic representation memory is empty, and no new regional coordinate system is established. The robot starts at position (a) in Fig.\u00a02. 2. After finding the landmark, the robot approaches to a certain distance in front of the landmark. For landmark recognition and distance estimation, we used a single consumergrade camera sensor to record and register the distances between the landmarks and the camera in advance of the robots motion. We assumed that the registered distances and angles could be the ground truth. The distance from the camera to a landmark was determined by estimating the geometric relation between previously registered features and currently matched features [25]. In this paper, we assumed that the distance was accurately represented when the distance to the landmark observed by the robot was within a certain percentage of the first registration for the distance. This information was then used to create a regional coordinate system. Position (b) in Fig.\u00a02 shows that the robot is approaching landmark for observation. We assume that the robot has a function that enables it to approach the landmark from the front and at a certain distance, and is able to measure using an landmark recognition system with a single camera when landmarks are in front of the robot and at a fixed distance. We also assume that the robot is able to move to this position within a certain tolerance of errors according to the results for the recognized landmark, estimated distance, and angle. 3. Position (c) in Fig.\u00a02 represents the creation of a regional coordinate system. If the recognized landmark did not be in a previous memory, a new node is established and a regional coordinate system is set up. If it is node , the regional coordinate system 1 3 is represented by \u03a9 = \u27e8 v, \u27e9 . The representation of the regional coordinate system creates the spatial landmark relations as follows. 4. In this case, the robot makes spatial landmark relations about the recognized landmarks around it. If no other landmarks are observed, then the semantic mapping is complete with the current node" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002434_s1068798x18080087-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002434_s1068798x18080087-Figure1-1.png", + "caption": "Fig. 1. Input shaft of a cylindrical gear [1].", + "texts": [ + "eywords: radial clearance, contact angle, contact deformation, radial and axial dynamic loads DOI: 10.3103/S1068798X18080087 A radial ball bearing of 0000 type may be used to withstand purely axial loads. As an illustration, we show the input shaft of a cylindrical gear in Fig. 1 (from [1]); the input shaft of a bevel gear in Fig. 2 (from [2]); and the supporting roller of a four-high stand in a high-speed rolling mill in Fig. 3 (from [3]). Each shaft is mounted on two radial roller bearings intended to withstand the radial load and one radial ball bearing absorbing the purely axial load. This is ensured by a hole drilled in the housing under the bearing, whose diameter exceeds that of the bearing\u2019s external race. The radial ball bearing is used in horizontal and vertical electrical machines, even with relatively large axial loads on the shaft created by the drive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003630_glocom.2018.8647980-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003630_glocom.2018.8647980-Figure1-1.png", + "caption": "Fig. 1. Conceptual illustration of the cross-scale TDD system,", + "texts": [ + " INTRODUCTION Biodegradable semiconductor materials [1] and engineered bacteria[2], combined with nanorobotic technologies [3], may find important applications in the medical field. These inorganic or organic miniature robots can physically disappear in the human body after completing the required tasks and cause no harm. Motivated by these emerging technologies, we have proposed the touchable communication (TouchCom) framework to investigate the performance of a cross-scale transient nanorobotic platform for targeted drug delivery (TDD) [4],[5]. As shown in Fig. 1, this cross-scale TDD system includes an external macro-unit (MAU) and a number of in vivo, drug-loaded micro-units (MIUs), which may be in the form of biodegradable silicon-based electronics [1] or engineered bacteria [3]. The MAU directs the motion of a swarm of nanobots by generating a guiding field [3]-[6]. For tracking of the swarm, drug particles can be labeled with fluorescent or contrast agents for imaging purpose and enables the drug cargo to feed the location of the swarm directly back to the MAU [7], [8]. TouchCom provides a one-to-one correspondence between TDD and communication systems as also illustrated in Fig. 1, which allows for analysis of the nanobots-assisted drug delivery process from the information transmission perspective. In our previous work [9], we have proposed a multi-physics channel model to describe the propagation of nanobots from the injection side to the target area. This model is comprised of three consecutive components as shown in Fig. 2, which are a fractal-based angiographically resolvable region, fractal-based angiographically unresolvable region, and near-target region. The first two components are called the far region, which corresponds to the arteries and is modeled using a fractal-based bifurcating structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003292_gtsd.2018.8595540-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003292_gtsd.2018.8595540-Figure1-1.png", + "caption": "Figure 1. Aging wine mechanism", + "texts": [ + " When supplying voltage to 220V into the plasma chamber where the voltage is amplified up many times from (0-220V) up to (20-50KV). With the high voltage between the negative and positive poles producing shock waves, ultraviolet rays, and breaking chemical bonds of the wine flow through the processing zone. After treated, alcohol is provided directly to the user through the tap. From experimental testing, the machine reached stability, wine output after treatment will meet the requirements of the Ministry of Health. Applying plasma technology helps break chemical bonds by using higher oxidants and electron dynamics. Figure.1 shows the aging wine mechanism. The plasma reactor is a coaxial cylindrical DBD cup includes two elecroders (out side cathode and inner anode) and a dielectric. Wine is filled inside a cup. A plasma source with high frequency and high voltage signal is applied to the two electrodes for a time sufficient to crease very strong oxidants such as HO*, O*, H*, O3, and H2O2. Simultaneously, the system produce a shockwave (dynamic electron and ions), ultraviolet (UV) in the N. Dam Tran, Nob. Harada Developing of Aging Wine Model by Dielectric Barrier Discharge (DBD) at Atmospheric Pressure 2018 4 th International Conference on Green Technology and Sustainable Development (GTSD) compound" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000619_6.2008-493-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000619_6.2008-493-Figure2-1.png", + "caption": "Figure 2: Four views of the final Hyperion design", + "texts": [ + " Four views of the Hyperion can be seen in . The Hyperion will take-off in less time than a fighter aircraft, when scrambled. It will then accelerate to an optimal climb speed, after which an optimal climb is initiated in the direction of the aircraft to be intercepted. At cruising altitude the Hyperion will level out, accelerate to maximum cruise speed and intercept the aircraft. Once arrived at the aircraft it will perform its mission, observation or escort, and subsequently land by using the instrument landing system (ILS). Figure 2 Figure 2 VI. Layout and systems After performing investigations into the aerodynamic and stability characteristics of the Hyperion, a final wing plan and control surface sizing is found, the parameter values in Table 2 are known and Hyperion\u2019s planform, as is shown in , is determined. The Hyperion must be able to fly autonomously and be remote controlled and for this it has to have advanced avionic systems on board, which are mostly commercial of the shelf (COTS) products and are located in modules in the front of the fuselage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003909_ecace.2019.8679324-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003909_ecace.2019.8679324-Figure2-1.png", + "caption": "Figure 2: Definition of force, moments and velocity in body fixed coordinate.", + "texts": [ + " The remaining of the paper is arranged as follows: section 2 represents system modelling, section 3 describes controller design, section 4 evaluates the performance of the proposed controller and section 5 concludes the paper. The aim of this section is to represent the generalized modeling of pitch angle regulation of aircraft. The system is modeled by linearized model of aircraft with considering different flight condition parameters. Fig. 1 presents the illustration of pitch control system. Here, Xd, Yd and Zd present the aerodynamic force. The parameters \u03c3, \u03d5 and \u03b3a present the pitch, roll and elevator deflection angle respectively. Fig. 2 presents the force, moment and velocity components. Here, L, D and E present the force, moments and velocity element respectively. The angular rates are shown by p, q and r. The parameters p, q and r represent the roll, pitch and yaw axis. The velocity of roll, pitch and yaw axis can be represented by u, v and w respectively. The system modelling data is used from: NAV IONa [1]. The parameters of dimensional derivative Q = 35.80lb/ft2, QSc\u0304 = 38495ft.lb and (c\u0304/2uf) = 0.015 are used for this paper. Table 1 shows the longitudinal directional stability parameters. Assuming the aircraft can maintain uniform altitude and velocity. Therefore, the thrust force and drag force eliminates each other. Besides, The weight and lift correspondence each other. Considering the pitch angle remains unchanged. From Fig. 1 and Fig. 2, the rigid body equation can be obtained. The lateral directional movement has been defined as X \u2212mgS\u03c3S\u03c3 = m(u\u0307 + qv \u2212 rv) (1) D = \u2212Iy q\u0307 + rq(Iy \u2212 Ix) + Ixz(p 2 \u2212 r2) (2) The statement (1), (2) and (3) are nonlinear. The disturbance theory has been used to linearize the equations. Solving the aircraft problem, it is required to consider the following assumption as: rolling rate p = \u03d5\u0307 \u2212 \u03a5\u0307S\u03c3 , yawing rate q = \u03c3\u0307C\u03d5+\u03a5\u0307C\u03c3S\u03d5, pitching rate r = \u03a5\u0307C\u03c3C\u03d5\u2212\u03c3\u0307S\u03d5 , pitch angle \u03c3\u0307 = qC\u03d5 \u2212 rS\u03d5, roll angle \u03d5\u0307 = p+ qS\u03d5T\u03c3 + rC\u03d5T\u03c3 and yaw angle \u03a5\u0307 = (qS\u03d5+rC\u03d5)sec\u03c3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002353_icphm.2018.8448958-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002353_icphm.2018.8448958-Figure1-1.png", + "caption": "Fig. 1. Diagrammatic Sketch of the Axle-box Vibration Signal in a Running High-speed Train.", + "texts": [ + " Moreover, the temperature is still used as a heart of 978-1-5386-1164-7/18/$31.00 \u00a92018 IEEE warning parameter in these systems. Once the alarm is triggered, vibration acceleration data would be used to trace the reason and confirm the alarm further. The internal vibration signals of axle-box mainly will be affected by numerous external factors, such as the impact of track, the running speed fluctuation, the vehicle body vibration and electromagnetic interference, limited by the installation position of acceleration sensors, shown as Fig. 1. Therefore, to develop effective highspeed axle bearing online vibration monitoring and detection systems, the difficulties relating to feature extraction and threshold setting need to be resolved. In general, the processing methods for axle bearing vibration data can be divided into two major categories. The first one is to extract the fault characteristic frequencies and diagnose faults by various means of signal analysis. Numerous studies have made advances in this area [3-9]. Mishra et al. [4] proposed an approach to diagnosis defects in the rolling element bearing with the short-time angle synchronous averaged signal, using order tracking on the envelope of wavelet de-noised evaluation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002103_ccdc.2018.8408202-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002103_ccdc.2018.8408202-Figure2-1.png", + "caption": "Fig 2. The geometry relationship of UAV speed", + "texts": [ + " The relative distance of adjacent UAVs is shortened with the decrease of the flight speed. Base on formula (10) and (11), the elastic distance vector \u02c6 g id can be written as cos cos sin sin cos \u02c6 cos sin cos sin sin sin cos x y z i i i i i i i i g x y zi i i i i i i i i i d x z i i i i d d d vd d d d v d d \u03c3 \u03d5 \u03d5 \u03c3 \u03d5 \u03c3 \u03d5 \u03d5 \u03c3 \u03d5 \u03c3 \u03c3 \u2212 + = + + \u2212 + (12) The 30th Chinese Control and Decision Conference (2018 CCDC) 6107 Here, \u02c6 g id can be further simplified. Supposing all the UAVs in the formation are flying in a stable cruise mode, the velocity vector can be decomposed as shown in Fig. 2. In Fig. 2, xyv is the component of the speed vector v in the Oxy plane and , ,x y zv v v are the components of the speed vector on three axes in the inertial coordinate system, respectively. When the UAV is in cruise flight, \u02c6 gd in (12) can be rewritten as \u02c6 g i i i id DW Z= + (13) Where, cos cos0 0 0 , cos sin , 0 0 0 sin x y i ii i y xi i i i i i i i d x z i i i d d v D d d W Z v d d \u03c3 \u03d5 \u03c3 \u03d5 \u03c3 \u2212 = = = \u2212 In addition, iD denotes the elastic coefficient matrix and changes with the speed of the ith UAV. Therefore, the elastic distance vector \u02c6 g id makes the safe distance adaptive to the leader speed, which can avoid the collision effectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002149_s40032-018-0480-4-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002149_s40032-018-0480-4-Figure2-1.png", + "caption": "Fig. 2 Detailed dimensions of a blade and b disc", + "texts": [ + " Disc Interface In the present work fir tree based disc and blade geometry was modelled with specific dimensions available from the literature [4, 7, 13]. Three geometric models were considered by varying contact angles at three levels, that is, 15 , 17.5 and 20 . Other geometrical parameters such as top flank angle = 40 , bottom flank angle = 40 , flank length = 3.556 mm, inner radius of the teeth = 0.5207 mm, outer radius of the teeth = 0.6833 mm, inner radius of the disc = 31.75 mm, outer radius of the disc = 190.5 mm, thickness = 20 mm and number of teeth = 3 were considered. Detailed dimensions of blade and disc is as shown in the Fig. 2. Inconel 720 nickel alloy was selected because of its excellent high temperature strength, oxidation and corrosion resistance [14]. The following properties of the material, that is, Young\u2019s modulus (E = 220 GPa), poisons ratio (t = 0.29), density (q = 8510 kg/m3) were considered.The meshed model of the blade disc interface is as shown in the Fig. 3a. SOLID 186, a higher order 3D, 20-node solid element which exhibits quadratic displacement behavior was considered for meshing purpose. Also some of the patch confirming tetrahedron elements were adopted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000294_14399776.2008.10785987-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000294_14399776.2008.10785987-Figure8-1.png", + "caption": "Fig. 8: Sketch of a pneumostatic bearing", + "texts": [ + " 6 and 7 show pressures along the air gap measured for the annular orifice supply system and simple orifice with feed pocket respectively, with relative supply pressure pS = 0.5 MPa and air gap heights h of 9 and 14 \u03bcm. Fig. 5: Photo of the air pad test bed For these systems, pressure distribution and mass flow rate measurements were used to calculate hole discharge coefficients Cd as a function of air gap height and supply pressure. The aerostatic journal bearing for linear motion constructed is shown schematically in Fig. 8. (1) is the journal, (2) is the bearing of diameter D = 50 mm and axial length L = 60 mm, provided with four equally spaced air inlet holes of diameter dB = 0.4 mm. The radial air gap is designated as h; and when journal and bearing are coaxial, h = h0 = 20 \u03bcm. Each hole communicates with a pocket of axial length Lp = 44 mm, width a = 15 mm, and depth \u03b4 = 100 \u03bcm. Bearing mass is 1.6 kg. An application of this type of bearing is shown in the photograph of Fig. 9. 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 8: 20 3 0 Se pt em be r 20 15 48 International Journal of Fluid Power 9 (2008) No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000687_ichr.2009.5379565-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000687_ichr.2009.5379565-Figure1-1.png", + "caption": "Fig. 1. Angles and frame s definition", + "texts": [ + " The last section of this paper is devoted to conclusions and perspectives . RABBIT IN ABSOLUTE COORDINATES The planar walking robot named Rabbit has five bod ies and 7 DOF in the sagittal plan. It has a trunk, two legs without feet. The angular coordinates referenced by the vertical are represented on Fig.l. The cartesian coordinates are referenced by (x ,z), coordinates of the hip in the fixed frame. During the gait (walking, running), the stance leg is in contact with the ground at a fixed point and has a rotational movement on this point. On figure 1, the right leg (coordinates ql and q2) is supposed to touch the ground and the left leg (coordinates q4 and q5) is in swinging phase and behind the hip. The dynamic model of the robot is given by the following equation : D(q)q+H(q,q)-Q(q) = Br+Ac1 (q)Fr+Ac2(q)FI (1) where q = [ql q2 q3 q4 q5 x zf is the absolute coordinate vector, q = [tiI q2 q3 q4 q5 X i ]T is the speed vector, D(q) is the mass matrix, H(q,q) is the centrifugal and Coriolis forces vector, r the torque vector on the joint (positive direction is indicated on Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003530_978-3-030-11220-2_34-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003530_978-3-030-11220-2_34-Figure4-1.png", + "caption": "Fig. 4. Example of two-stage gearbox operating in underground mine.", + "texts": [ + " how many values one appears in the vector. We propose to classify machine state based on following principles: \u2013 If the value is smaller than 5% machine is healthy \u2013 If the value is between 5 to 50% machine is in the warning state \u2013 If the value is greater than 50 % machine is damaged. The flowchart of the procedure is presented in Fig. 3. In this section we analyze three signals comes from the real life gearbox. It is a part of a belt conveyor which is operating in underground mining industry. In Fig. 4 we present the measurement sensors placement (top panel) and example gearbox operating in underground mine (bottom panel). It is a two-stage gearbox with first stage being conical, and second \u2013 cylindrical. Simplified kinematic schematic of that type of gearbox is presented in Fig. 5. Frequency parameters of the gearbox are presented in Table 1. In the analyzed vibration data the sampling frequency is 17066 Hz. All signals have 1022676 observations, which translates to almost 60 s. Figure 6 (left) shows raw vibration data of all 60 s corresponding to healthy (top panel), in warning state (middle panel) and damaged (bottom panel) machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001529_s12008-018-0471-y-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001529_s12008-018-0471-y-Figure7-1.png", + "caption": "Fig. 7 The 3D finite element model", + "texts": [ + " The mesh size of liner is 0.25 mm. The mesh size of load plate and mandrel is 1 mm. The total number of nodes is 383447 and the elements are 93832. The analysis element is solid 185, and this element is defined by 8 nodes having three degrees of freedom at each node. The boundary conditions are as follows: the both ends of mandrel are fixed completely; the outer ring surface of the load plate is set to full constraint; the radial load of 44KN is applied to the load plate. The finite element model is shown in Fig. 7: The simulation results are shown in Fig. 8. The simulation results from Fig. 8 showed that the maximum displacement between the liner and inner ring occurred on the axial cross-section, this section is the intersectionplane between XOY planes. Thence, the maximum wear depth is on the same section. With the intention of simplify the calculation and the debugging of the program, the middle section where themaximumdisplacement between the inner ring and the outer ring of the self-lubricating joint bearing is taken as an object" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001099_icici-bme.2009.5417228-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001099_icici-bme.2009.5417228-Figure3-1.png", + "caption": "Fig. 3. Exemplary implementation of a health monitoring system on a wheelchair, providing localization data via GPS, but also crucial health parameter in order to organize quick support an emergency.", + "texts": [], + "surrounding_texts": [ + "Mobility is often constrained by fears of the handicapped persons that they might encounter problems, while being on the way. Therefore modern telecommunication and data \u00b4processing techniques are provided to establish a quick information of health care centers, in case an emergency occurs. This way patients can be monitored at their homes as well as travelling to receive immediate feedback and related help [6]. Typical equipment includes navigation instruments (in order to send an emergency team directly to the appropriate location, as well as sensors appropriate to the specific handicap to characterize the health status (in order to provide the emergency team a decision basis which medicine they should take with them before they depart from the health care center) [4]. In several areas (cardiology, nephrology,\u2026) such telemedicine applications have already been evaluated in clinical tests and proved significant cost reductions due to the fast detection of evolving changes in the health status. Despite investments in the telematics infrastructure nevertheless an overall reduction of costs to more than 50 % could be reported [1]. The information chain is depicted in Fig.4 and extends as follows: Anhand der gew\u00e4hlten Problemstellung von \u201eFernbetreuung bei COPD und Heimbeatmung\u201c kann die gesamte Informationskette entwickelt werden: patient --> medical intruments --> data acquisition and preprocessing --> data transmission --> heaalth center with specialists decide about interpretation of data and needed activities --> potential activities: a) call and consultancy of patient, b) information of the responsible doctor of the patient, c) initiation of emergency activities." + ] + }, + { + "image_filename": "designv11_92_0002489_s13198-018-0747-4-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002489_s13198-018-0747-4-Figure2-1.png", + "caption": "Fig. 2 The 3D models of equilibrium elbow", + "texts": [ + " 39: w1 \u00bc fk Xl fk Xk f1 Xk\u00f0 \u00de f1 Xl\u00f0 \u00de\u00bd \u00fe f2 X k 1\u00f0 \u00de\u00f0 \u00de f2 X2\u00f0 \u00de\u00bd \u00fe \u00fe fk Xl\u00f0 \u00de fk Xk\u00f0 \u00de\u00bd w2 \u00bc fk 1 X2\u00f0 \u00de fk 1 X k 1\u00f0 \u00de f1 Xk\u00f0 \u00de f1 Xl\u00f0 \u00de\u00bd \u00fe f2 X k 1\u00f0 \u00de\u00f0 \u00de f2 X2\u00f0 \u00de\u00bd \u00fe \u00fe fk Xl\u00f0 \u00de fk Xk\u00f0 \u00de\u00bd .. . wk \u00bc f1 Xk f1 X1\u00f0 \u00de f1 Xk\u00f0 \u00de f1 Xl\u00f0 \u00de\u00bd \u00fe f2 X k 1\u00f0 \u00de\u00f0 \u00de f2 X2\u00f0 \u00de\u00bd \u00fe \u00fe fk Xl\u00f0 \u00de fk Xk\u00f0 \u00de\u00bd 8 >>>>>>< >>>>>>: \u00f039\u00de We propose two optimal objectives: one is the whole mass of the component lighter and the other is the reliability sensitivity of the structure parameters can be decreased to the minimum which make the equilibrium elbow having the robust characteristic. The basic size of the equilibrium elbow as shown in Fig. 2 is a \u00bc 60; 1:2\u00f0 \u00demm, b \u00bc 35; 1:8\u00f0 \u00demm, c \u00bc 30; 2:5\u00f0 \u00demm, r1 \u00bc 45; 0:8\u00f0 \u00demm, r2 \u00bc 45; 1:0\u00f0 \u00demm, r3 \u00bc 10; 0:08\u00f0 \u00demm, the material strength of the equivalent bow is r \u00bc 145; 5:125\u00f0 \u00deMPa, the torque loading of the equilibrium elbow are loading variables following arbitrary life distribution, and the fourth moment is: T \u00bc 3:2074 106N; 1:5791 105N2; 1:2189 1015N3; 2:0153 1021N4\u00de. Setting the structural parameters and the loading as the random variance to calculate the reliability sensitivity, Xs \u00bc \u00bdr1; r2; r3; a; b;M; T T " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003990_j.enganabound.2019.04.008-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003990_j.enganabound.2019.04.008-Figure5-1.png", + "caption": "Fig. 5. Sketch map of finite line contact pair.", + "texts": [ + " When the boundaries have been dispersed, the ntegral equation of each object is expressed as follows: \ud835\udf091 \ud835\udc51 \ud835\udf092 I { 4 \u2211 \ud835\udefd=} + } w \u0394 i d r t i m r n t n 3 3 r a l w \ud835\udf0c t w t o w r o p e w u b c w r \ud835\udc36 \ud835\udc58 \ud835\udc56\ud835\udc57 ( \ud835\udc4b )\u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a \ud835\udc57 ( \ud835\udc4b ) + \ud835\udc41 \ud835\udc58\ud835\udc53 \u2211 \ud835\udc5b =1 { \ud835\udc5e \u2211 \ud835\udefd=1 \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc47 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d = \ud835\udc41 \ud835\udc58\ud835\udc53 \u2211 \ud835\udc5b =1 { \ud835\udc5e \u2211 \ud835\udefd=1 \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d \ud835\udf091 \ud835\udc51 \ud835\udf092 } + \ud835\udc41 \ud835\udc58\ud835\udc4f\u2211 \ud835\udc5b =1 + \ud835\udc41 \ud835\udc58\ud835\udc4f I \u2211 \ud835\udc5b =1 { 4 \u2211 \ud835\udefd=3 2 \u2211 \ud835\udc59=1 \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf09\ud835\udc59 \ud835\udefc\ud835\udc59\ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d \ud835\udf091 \ud835\udc51 \ud835\udf092 + \ud835\udc41 \ud835\udc58\ud835\udc4f II \u2211 \ud835\udc5b =1 { 2 \u2211 \ud835\udefd=1 2 \u2211 \ud835\udc59=1 \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf09\ud835\udc59 \ud835\udefc\ud835\udc59\ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d \ud835\udf091 \ud835\udc51 \ud835\udf092 here, q represents the node number of one element, \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udc57 and \ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udc57 represent incremental displacement and traction of the \ud835\udefdth node n the n th element of k -object in the j direction under the global coor- inate system respectively when the m th step increment is loaded, \ud835\udefchj epresents the direction cosine between the local coordinate system and he global coordinate system, \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf09\u210e and \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf09\u210e \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udc57 represent ncremental displacement and traction of the \ud835\udefdth node in the n th ele- ent of k -object in the direction of \ud835\udf09h under the local coordinate system espectively when the m th step increment is loaded, N kf represents the umber of non-contact elements of k -object, N kb I and N kb II represent he numbers of BBE1 and BBE2 of k -object, and N kc represents the total umber of contact elements of k -object. . EHL theory of finite length line contact .1. Mathematical model Finite length line contact pair is composed of two end modification oller A and B, as shown in Fig. 5 . Where, R x and R y represent the radius nd the repair radius of the rollers, L and l represent the length and the ine segment length of rollers [19] . (1) Reynolds equation of finite length line contact is expressed as follows: \ud835\udf15 \ud835\udf15\ud835\udc65 ( \ud835\udf0c\u210e 3 \ud835\udf02 \ud835\udf15\ud835\udc5d \ud835\udf15\ud835\udc65 ) + \ud835\udf15 \ud835\udf15\ud835\udc66 ( \ud835\udf0c\u210e 3 \ud835\udf02 \ud835\udf15\ud835\udc5d \ud835\udf15\ud835\udc66 ) = 12 \ud835\udc62 \ud835\udc60 \ud835\udf15( \ud835\udf0c\u210e ) \ud835\udf15\ud835\udc65 (9) here, p represents the oil film pressure, h represents the film thickness, represents the oil density, \ud835\udf02 represents the oil viscosity, u s represents he suction velocity. (2) Boundary conditions equations are expressed as follows: { \ud835\udc5d ( \ud835\udc65 \ud835\udc56\ud835\udc5b , \ud835\udc66 ) = \ud835\udc5d ( \ud835\udc65 \ud835\udc5c\ud835\udc62\ud835\udc61 , \ud835\udc66 ) = \ud835\udc5d ( \ud835\udc65, \ud835\udc66 \ud835\udc56\ud835\udc5b ) = \ud835\udc5d ( \ud835\udc65, \ud835\udc66 \ud835\udc5c\ud835\udc62\ud835\udc61 ) \ud835\udc5d ( \ud835\udc65, \ud835\udc66 ) \u2265 0 ( \ud835\udc65 \ud835\udc56\ud835\udc5b < \ud835\udc65 < \ud835\udc65 \ud835\udc5c\ud835\udc62\ud835\udc61 , \ud835\udc66 \ud835\udc56\ud835\udc5b < \ud835\udc66 < \ud835\udc66 \ud835\udc5c\ud835\udc62\ud835\udc61 ) (10) w r u w w ( c } + \ud835\udc41 \ud835\udc58\ud835\udc50 \u2211 \ud835\udc5b =1 { \ud835\udc5e \u2211 \ud835\udefd=1 \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf09\ud835\udc59 \ud835\udefc\ud835\udc59\ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc47 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d \ud835\udf091 \ud835\udc51 \ud835\udf092 } 3 \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf093 \ud835\udefc3 \ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd I ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d \ud835\udf091 \ud835\udc51 \ud835\udf092 } \ud835\udc41 \ud835\udc58\ud835\udc4f II \u2211 \ud835\udc5b =1 { 2 \u2211 \ud835\udefd=1 \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a,\ud835\udc5b,\ud835\udefd \ud835\udf093 \ud835\udefc3 \ud835\udc57 \u222c\u0393\ud835\udc5b \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 [ \ud835\udc4b, \ud835\udc4c ( \ud835\udf091 , \ud835\udf092 ) ] \ud835\udc41 \ud835\udefd II ( \ud835\udf091 , \ud835\udf092 ) ||\ud835\udc3a( \ud835\udf091 , \ud835\udf092 ) ||d \ud835\udf091 \ud835\udc51 \ud835\udf092 } (8) here, x in and x out represent the entrance and exit coordinate of x direcion, respectively, y in and y out represent the entrance and exit coordinate f y direction, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003769_1.5092451-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003769_1.5092451-Figure2-1.png", + "caption": "Fig. 2. Material needed for the diamagnetic levitation experiment.", + "texts": [ + " 271 displace the magnet from the equilibrium point it undergoes oscillations around the equilibrium. These oscillations are damped by friction. During the second experimental session, we provided the students with the following material: a set of many diskshaped neodymium magnets, a metallic (iron) piece with an angular shape making a 90 degree angle and 120 mm in length, two aluminum supports, a small compass-like device that helps to determine magnetic poles, and a graphite pencil mine 60 mm long (see Fig. 2). The iron piece is intended to aid the students to position the magnets. With all of these elements, the students are asked to find a configuration in which the graphite mine levitates freely.7 All of the groups of students have been able to achieve levitation and they found several combinations in which a stable levitation is possible. Figure 3 shows a typical result. Superconductors behave as superdiamagnetic materials and can be seen as magnetic mirrors: the magnetic field generated by the currents in the superconductor can be described as having been originated by a magnet equal in magnitude to the external magnet, and with like poles opposing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001617_s40799-018-0232-7-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001617_s40799-018-0232-7-Figure3-1.png", + "caption": "Fig. 3 Testing bench section in ANSYS\u00ae Academic with loads,", + "texts": [ + " Once the training is finished, the students have covered the topics related to drawing geometries, assignment of materials, discretisation of models, symmetry, sub-modelling, types of contact, non-linearities, parameterization and data processing. The whole analysis methodology involves: firstly, the geometry of the testbed should be imported to the geometrical module of ANSYS \u00ae 16.2. Then, it is necessary to introduce the steel properties listed in Table 1. Next, the model should be discretised and its constraints should be properly defined. Finally, a parameterized load is defined at the frontal ring, also a rectangular section, which represents a strain gage being used at the validation stage, is suggested to be defined as seen in Fig. 3. Specifically, the model shown in Fig. 3 considered a fixed support (ux = uy = uz = 0) on the far end of the circular hollow bar, tagged with letter A; also, initial element size in this model is 4 mm, but students should be encouraged to try various element sizes. It is proposed to progressively add equivalent loads by adding masses from 0.1 to 1 kg at the ring of the solid cylinder in order to solve the finite element model with iterations of different loads. The initial results are the longitudinal and tangential strains, at the defined area which represents the attached strain gage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002393_gt2018-75983-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002393_gt2018-75983-Figure1-1.png", + "caption": "Figure 1 Schematic of single-pad HAFB", + "texts": [ + " The bump foil strip is constructed with 33 bumps, distributed circumferentially, and with a constant bump height along the axial direction. Both the top foil and bump foil are made of Inconel X750 (0.127 mm thickness). The bearing sleeve has seven holes for orifices (six on the top and one on the bottom), which allows for the configuring of HAFBs with different hydrostatic injection locations. The circumferential locations of the orifice holes on the bearing sleeve are 30, 60, 90 180, 270, 300, and 330 degrees (as shown in Figure 1). In the axial direction, orifice holes are located in the middle of the bearing sleeve. Figure 2 shows the single-pad HAFB. Overall parameters and geometries of the test HAFBs are presented in Table 2. The externally pressurized air is injected into the HAFB radial clearance through orifice tubes (stainless steel) welded onto the backside of the top foil. Three sets of single-pad HAFBs are constructed with different circumferential positions of orifices. The first set of HAFBs (case-1) has orifices located at 30, 180, and 330 degrees (Figure 1). The orifices of the second set (case-2) are located at 60, 180, and 300 degrees. For the third set (case-3), the orifices are located at 90, 180, and 270 degrees. In general, all three cases have a single orifice at the bottom (opposite to the loading direction), but the angular locations of the two top orifices are different for each case. Figure 1 shows a schematic view of the case-1 single-pad HAFB. The orifice configuration the three different sets is summarized in Table 1. 3 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 09/17/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Rotor consists of five main components: two bearing journals, an induction motor element, a thrust runner, and an end cap (with the same weight as the thrust runner). All the components are assembled with a set of tension bolts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003231_icsai.2018.8599363-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003231_icsai.2018.8599363-Figure3-1.png", + "caption": "Figure 3. The autonomous air aircraft carrier\u2019s airborne pod structure model that grips the pull rod on the fuselage of UAV, the UAVs system are hung on the wing of the large passenger plane.", + "texts": [ + " When griping the pull rod under the belly of UAV, the flower gripper structure of the retractable connecting rod column firmly seizes the pull rod of the landing rod of an UAV positioning UAV in the location of the slab on the taxiing runway. When the retractable rod column of slab of air aircraft carrier shrinks to its minimum, an UAV is safely positioned on air aircraft carrier\u2019s slab. The rod is able to stretch to large passenger aircraft slide the front end of the runway. The root of the rod column (tie bar) is located in the front of the aircraft cockpit, as shown in Figure 2. In Figure 3, the model of pod and pod mounting system of the airborne that grips the pull rod on the fuselage of UAV is showed. When there are flight accidents of large aircraft on geographic region, or when large aircraft has strong oscillation and vibration, it releases a UAV on top of flight, and the UAV takes off on the taxiing runway on top. Or, an autonomous aircraft carrier (AAC) release its pod-UAV system under its wing. The UAV can accompany with flying of large passenger aircraft in the sky, and the UAV monitors and tracks large passenger aircraft flight, and passes the large aircraft flight state and routes, and starts the internal software robot (software UAV) that provides the internal state and video information of large passenger aircraft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001482_jqme-01-2016-0003-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001482_jqme-01-2016-0003-Figure8-1.png", + "caption": "Figure 8. Interaction of a healthy gear driving a gear with a tangential root crack", + "texts": [ + " The sensor was mounted in a housing such that it was axially aligned with the shaft/magnet and positioned approximately 2-4 mm away in the axial direction (Figures 6 and 7). The sensor output is a voltage that varies sinusoidally (one cycle per rotation of the shaft). The gears were placed on the shaft such that the direction of rotation for the drive motor was appropriate for the fault being tested. This direction was irrelevant for the radial root crack but was crucial for the tangential root crack. The direction of rotation for the tangential root crack coincided with the direction of crack growth. Figure 8(a) illustrates the running direction for the driving gear with the presence of a tangential root fault shown. With this direction of rotation and torsional loading, the corresponding driving forces try to force the crack open causing the fault to propagate in the manner illustrated. The output from the Hall Effect sensors is a sinusoidal curve relative to the absolute angular displacement between the sensor and the magnet/shaft. The dynamic range in this case was approximately \u00b11.7 V DC due to the sensor placement relative to the magnet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003164_j.ifacol.2018.11.538-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003164_j.ifacol.2018.11.538-Figure3-1.png", + "caption": "Fig. 3. The developed mini-drone", + "texts": [ + " In addition, an RC receiver, FM800, is connected to the microcontroller, and it is utilized during manual flight and to switch flight modes. Figure 1 shows a connection diagram of the components described above. The PCB design is completed by using EAGLE CAD software, and all PCBs are designed with four layers. The main PCB wheelbase is 130 mm and the dimensions of the sensor PCB is 35 \u00d7 35 mm. As was mentioned in Chapter 2.1, the ground plane is incorporated into the sensor PCB to improve GNSS signal reception. Figure 2 shows the PCB design results (left) and the manufactured sample PCBs (right); Fig. 3 shows the completed mini-drone, with all electronic parts soldered onto the PCBs and several parts installed, including the propeller, coreless motor, and RC receivers. The developed mini-drones are capable of autonomous flight in an outdoor environment; the detailed specifications are summarized in Table 2. IFAC SYROCO 2018 Budapest, Hungary, August 27-30, 2018 180 Dasol Lee et al. / IFAC PapersOnLine 51-22 (2018) 178\u2013183 Kalman filters (Kalman, 1960), extended Kalman filters, and unscented Kalman filters (Julier, 2004) are widely utilized for drone navigation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.38-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.38-1.png", + "caption": "Fig. 11.38 Generalized skewed-parabola Dgsp(x1/l; n, f) describing the contact pressure profile in the circumferential direction. Reproduced from Ref. [10] with the permission of Tire Sci. Technol.", + "texts": [ + " Furthermore, the pressure distribution becomes asymmetric in that it is inclined forward. The pressure distribution given by Eq. (11.72) therefore becomes a generalized skewed-parabola function that expresses the forward inclination of qz(x1): qz\u00f0x1\u00de \u00bc n\u00fe 1 n Fz wl Dgsp x1 l ; n; f Dgsp x1 l ; n; f \u00bc 1 2 x1 l 1 n 1 f 2 x1 l 1 n o : \u00f011:145\u00de The generalized skewed-parabola function Dgsp(x1/l; n, f) can easily express the forward inclination of qz(x1) using the parameter n to control the shape of the pressure distribution and f to control the inclination as shown in Fig. 11.38. The fundamental equations and the procedure used to calculate the cornering properties of a tire in a combined slip (in braking) are as follows. (i) A generalized skewed-parabola function is used for the pressure distribution Dgsp(x1/l; n, f) according to Eq. (11.145). (ii) The pressure distribution in the circumferential direction receives feedback in the form of the inclination parameter f while the circumferential movement of the contact patch xc is calculated using either f \u00bc CMfMz xc=l \u00bc 1=2 CMcMz=l 2 \u00f011:146a\u00de or f \u00bc CfFx xc=l \u00bc 1=2\u00feCxcFx: \u00f011:146b\u00de (iii) Mz is fed back to the slip angle of the wheel a0 to calculate the effective slip angle a21: a \u00bc a0 Mz=Rmz: \u00f011:147\u00de (iv) The sliding point rh is determined (where rh = 0 is used in the case that rh < 0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000779_icit.2009.4939600-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000779_icit.2009.4939600-Figure1-1.png", + "caption": "Fig. 1. The structure in the wind end region of a flameproof motor.", + "texts": [ + " So the edge element method is extensively used in motor magnetic field analysis. As edge element method can not directly calculate the inductance of solid source conductor, it can not realize direct circuit field coupling. In this paper, a novel methodology for coupling edge element analysis with circuit simulation is presented. The distribution of eddy current field in the iron core and core loss is calculated. For a flameproof induction motor, the typical structures of the end region are shown in Fig. 1. The full 3-D model of the stator end windings are shown in Fig. 2. Fig. 3 shows the circuit diagram of the flameproof motor. To impose loads on the 3-D finite element model expediently, the circuit-field coupled approach is adopted. Two basic approaches to coupling finite element analysis with circuit simulation exist. One is direct coupling approach [5] where the field and circuit equations are coupled directly together and solved simultaneously. The other approach is indirect coupling where the finite element analysis and circuit simulator are treated as separate systems in a step-by-step process with respect to time, while they exchange coupling variables in each step [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003117_978-3-030-01382-0_6-Figure6.13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003117_978-3-030-01382-0_6-Figure6.13-1.png", + "caption": "Fig. 6.13 Front view of the aircraft roll motion in order to do horizontal turning", + "texts": [ + "12 describes the horizontal geometry of the problem. The radial component of the velocity vector V (the one that is pointing along the direction of R) is V cos(w \u2212 e). This velocity component acts to reduce R. The component of the velocity vector V that is vertical to R is V sin(w \u2212 e). This component acts to reduce e. The motion equations are therefore dR dt \u00bc V cos\u00f0w e\u00de \u00f06:8\u00de R de dt \u00bc V sin\u00f0w e\u00de \u00f06:9\u00de For the airplane to make a turn in a horizontal plane, it needs to roll, at an angle / about its body axis, as shown in Fig. 6.13. The reason for rolling the body of the aircraft is to create a lift force L in a direction which is not equal to the Earth gravity direction. One vector component of L balances the aircraft\u2019s weight. The other vector component of L creates a centripetal force that causes the airplane to steer in a horizontal plane. L cos/ \u00bc mg \u00f06:10\u00de L sin/ \u00bc mV2 r \u00bc mV dw dt \u00f06:11\u00de The negative sign in the second equation above comes from an arbitrary definition of the positive direction of the rotation /. Dividing the above equations by one another yields the relationship: dw dt \u00bc g V tan\u00f0/\u00de \u00f06:12\u00de It is interesting to note that dw/dt is independent of the lift force L" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003335_phm-chongqing.2018.00209-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003335_phm-chongqing.2018.00209-Figure5-1.png", + "caption": "Fig. 5. The location of the fault bearing, the fault gear and the sensor", + "texts": [ + " The bearing is N1007 (the number of rolling element Z=16, the contact Angle = 0, the diameter of rolling element d=6.75mm, and the diameter of pitch circle d= 47.5mm), in Fig. 4. The speed of Input shaft slowing down from 700 rev/min. The data acquisition card is NI USB9234, sampling frequency is 25.6 kHz, sampling time 50 s, charge amplifier is DH - 5853, the type of the acceleration sensor for DH112, sensitivity 5.20 pC/g. The position of acceleration sensor and eddy current displacement sensor is shown in Fig. 5. The fault feature order of N1007 bearing is computed as follows: 1 1(1 cos ) 6.88 2outer r r Z d f f f D \u03b1= \u2212 = (8) Where, fr1 is the rotation frequency of the bearing inner race. Therefore, the bearing fault feature order is 6.88X. Select input shaft as the reference axis, the order of gear fault feature is 1X . The original vibration signal as shown in Fig. 6, the original key-phasor signal (in order to display, this thesis only intercept 0 to 10 seconds, the actual calculation of 0 to 50 seconds) as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001350_20080706-5-kr-1001.01155-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001350_20080706-5-kr-1001.01155-Figure5-1.png", + "caption": "Fig. 5. Rotation of obstacle force.", + "texts": [ + " In other words, by turning the effect of obstacle forces towards the leader\u2019s direction. Say \u03b8OB is the direction of FOB and \u03b8V L is the direction of the virtual leader. We apply the vector field \u03a6 to the obstacle forces. The resulting force is and its direction is given in equation (19) and (21). FOBnew = \u03a6(FOB) (19) \u03a6(FOB) = |FOB |(icos\u03b8OBnew + jsin\u03b8OBnew ) (20) \u03b8OBnew = \u03b8OB + \u03b8V L 2 if \u03b8OB \u2212 \u03b8V L < \u03c0 \u03b8OB + \u03b8V L 2 + \u03c0 if \u03b8OB \u2212 \u03b8V L > \u03c0 (21) The rotation of obstacle force is illustrated in figure 5. This modification shows greater performance for cornerings. In corners when virtual leader turns sharp corners FV L and FOB acts on opposite directions for the vehicles in the back of the group. Since FV L is proportional to distance to leader and FOB is inversely proportional to distance to obstacle, vehicle in this figure ends up moving back and forth towards the obstacle surface. The rotation prevents this and makes the vehicle follow the leader more when an obstacle exists. Formation shape is given by desired intervehicle distances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003712_robio.2018.8665214-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003712_robio.2018.8665214-Figure7-1.png", + "caption": "Fig. 7. Waterborne test tracks (top view and side view)", + "texts": [ + " However, in the case of short spikes in low flow conditions, the wall surface and flow disturbances influence its low drag values. As described above, the long spikes generate sufficient thrust force to move the small crawler vehicle shown in Fig. 1 along the surface of still water. V. DEMONSTRATION TEST The tested spikes were attached to a radio-controlled crawler track, which was placed in a swimming pool for the demonstration test. The basic structure of the tested spiked track is similar to the grass-slope climbing prototype introduced in Fig. 1 and Table I. In this model, as shown in Fig. 7, the crawler dimensions are 30 cm in length and 30 cm width, with long spikes attached. The spikes are staggered, with alternating rows of three and four spikes. On the front and rear sides of the crawler, buoyant urethane floats are mounted. The total weight is 3 kg. The waterborne test was conducted in a 1.5-m deep swimming pool. The lower half of the crawler was completely submerged in the water, and rectilinear motion along with right and left turns were tested. The crawler vehicle began the test on the side of the swimming pool and subsequently dove toward the surface of the water" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002955_978-3-030-03451-1_54-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002955_978-3-030-03451-1_54-Figure1-1.png", + "caption": "Fig. 1. a. Main components of a permanent magnet rotor, b. 3D model of the \u201cProLemo\u201d rotor design, c. 3D model of the \u201cOptiMA\u201d rotor design", + "texts": [ + " Keywords: Assembly(ing) Modelling Optimization In the context of rising demands for electric vehicles, the production of high efficient permanent magnet synchronous motors gains more and more importance. Being a part of the production of electric motors, special focus needs to be put on deviations occurring during manufacturing and assembly of a rotor\u2019s components to ensure high quality standards at considerable costs. The rotor of a permanent magnet synchronous motor can be split up into four main components: shaft, balancing discs, rotor discs and magnets (Fig. 1a). Within this article, two rotor designs are used: Rotor discs for the ProLemo design are made of light-weight materials and are manufactured in a new injection moulding process (Fig. 1b). This injection moulding process has been developed within the \u201cProLemo\u201d project and corresponds to a process that is intended to offer an alternative to \u00a9 Springer Nature Switzerland AG 2019 R. Schmitt and G. Schuh (Eds.): WGP 2018, Advances in Production Research, pp. 551\u2013562, 2019. https://doi.org/10.1007/978-3-030-03451-1_54 lamination stacks [1]. Each rotor discs carries 12 magnets on its circumference. Rotor discs for the \u201cOptiMA\u201d design however are made of lamination stacks and 48 magnets are to be placed within each rotor disc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001145_ht2009-88365-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001145_ht2009-88365-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of sessile droplet and associated terminology.", + "texts": [ + " The VOF model has been previously used in related problems, such as in the prediction of droplet splashing [27,28]. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2009 by ASME In the present work, we address the prediction of the critical angle of inclination at which a droplet on an inclined plane begins to slide. A schematic diagram of the droplet and the associated terminology is shown in Figure 1. We employ the VOF-CSF model, with the contact angle variation along the contact line being specified as an input. The model employs advancing and receding contact angles obtained from experiment. The VOF simulation is then used to predict the critical angle of inclination at which sliding commences. The predictions of the critical angle of inclination so obtained are validated against experiments conducted as part of this work. In the current work the VOF-CSF model in FLUENT [29] is used. A custom contact-angle model is implemented using user defined functions to capture the azimuthal contact angle variation", + " For this purpose, Youngs\u2019 Geometric Reconstruction Scheme [30], which is based on piecewise-linear reconstruction of the interface in a partially filled cell, is used. Further details of the implementation may be found in the FLUENT manual [29]. The momentum equation is solved for the average velocity of the mixture, and the influence of multiple phases appears through the phase fraction-dependent local properties of the material in each cell. . Tv vv p v v g F t (3) Here, the gravitational acceleration is specified as cos sing g i g j , as shown in Figure 1. Physical properties such as density and viscosity are volume-averaged as follows (1 )s s s p (4) (1 )s s s p (5) where subscripts p and s represent the primary and secondary phases, respectively. In the CSF formulation, when only two phases exist, the volumetric force, F in (3) is given by: 1 2 s s p s F (6) where s is the interface curvature for the secondary phase, given by Brackbill et al. [31] to be: \u02c6.n (7) In (7), n\u0302 is the unit normal vector. The normal is obtained based on the volume fraction gradient given by: sn (8) The interface shape at the triple line, where the two phases meet the wall, is imposed by specifying n\u0302 through the specification of the contact angle as: \u02c6\u02c6 \u02c6 cos( ) sin( )w w w wn n t (9) where \u02c6wn and w\u0302t are the unit vectors normal and tangential to the wall directed into the secondary fluid and wall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002858_012090-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002858_012090-Figure1-1.png", + "caption": "Figure 1. 3D finite element model for the SLM of GH3536 powder.", + "texts": [ + " These physicochemical phenomena cannot be investigated by traditional experimental method, but finite 1234567890\u2018\u2019\u201c\u201d element method is a reasonable choice.8, 9 In the present work, a 3D finite element model is developed to study the effect of process parameters on temperature distribution and melting pool behavior during SLM. In addition, several GH3536 alloy samples are fabricated using SLM technology with different thermal input parameters and the thermal behavior is analyzed by experiments as well. Numerical simulation is performed using finite element modelling in this research. Figure 1 shows the geometry of the model with defined laser scanning path. The model consists of a stainless-steel substrate with a dimension of 1.40\u00d70.80\u00d70.20 mm, a powder bed of GH3536 with a dimension of 0.80\u00d70.40\u00d70.08 mm. The powder bed is divided into two layers. Each layer is 40 \u03bcm thickness. To reduce simulate time and maintain sufficient calculation accuracy, a finer mesh is used in the powder layer while coarser mesh is used in the stainless-steel substrate. The fine mesh size is 0.0125\u00d70.0125\u00d70" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000438_s1068371209080021-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000438_s1068371209080021-Figure4-1.png", + "caption": "Fig. 4. External view of drive prototype.", + "texts": [ + " The equation of motion at this interval of transient processes can be written as follows: (5) where T1 \u2013 is the period of resonance oscillations of the piezopipe 2 and holder 6 of the movable object. When we solve Eq. (5), the primary velocity and object displacement are determined according to Eq. (4) under t = t1. Equations (1)\u2013(5) make it possible to estimate the behavior of the transient processes in the inertial piezodrive when the sharp cutoff of the control signal is formed, as well as to choose a mode that allows one to simplify the application of electrodynamic system with the goal of improving drive performance. Testing of a drive model with an electromagnet (Fig. 4) shows the following: (1) If we use electromagnetic action on a movable holder, it is possible to increase the displacement accuracy for the concrete drive by about 18% (up to 20 nm). It is difficult to further increase the accuracy due to insufficient power of the electrical magnet, which in turn makes it impossible to completely compensate for the friction force. (2) Currents of up to 10 A can appear in the drive; therefore, it is necessary to remove the heat effectively T1 2d2x dt2 ------- 2\u03be K BL( )2 kRT1 ------------------+\u239d \u23a0 \u239b \u239e T1 dx dt ----- x+ + 0,= RUSSIAN ELECTRICAL ENGINEERING Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001121_2008-01-1408-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001121_2008-01-1408-Figure2-1.png", + "caption": "Figure 2: Tire deformation with a camber angle", + "texts": [], + "surrounding_texts": [ + "Copyright \u00a9 2008 SAE International\nABSTRACT\nBased on the simulation results of tire rolling over perpendicular cleats by MPTM model, in present paper, a series of simulation results of tire rolling over oblique cleats with different angles are given. For that, the Modal Parameter Tire Camber property Model is established. For the appraisement of comparison between simulation and experimental results a problem concern the validation test is pointed out. In the end, simulation results of tire rolling over a series of continuous cleats are given.\nINTRODUCTION\nSeveral high level analytical models are developed to meet the growing demand in automotive virtual proving ground simulation technology. They are SWIFT-Tyre [1], RMOD-K [2], FTIRE [3] and MPTM [4-12]. In paper [12], a series of simulation results of tire rolling over perpendicular cleats by MPTM model are given. They are quite agreed with the experimental results in both time and frequency domains. And a series of simulation capabilities of the model are presented. But for full virtual proving ground simulation, a 3-dimensional model should be developed. For example, simulation of tire rolling over oblique cleats is lateral force and moment response concern. For that, the modal parameter tire camber property model is established on the basis of the vertical model. Lateral modal parameters are introduced and tread width is considered. A steady state case calculated for a tire 205/55R16 with camber angle and fixed rim center in horizontal position is given. Vertical force is roughly verified by experiment in range of vertical deformation to 0.025m. Vertical, cornering and camber model are synthesized, and a three dimensional MPTM is formed. Simulations of a tire 205/55R16 rolling over oblique cleats with different incline angles are performed. A series of calculation results, such as: vertical, longitudinal, lateral forces responses; moment responses about vertical, longitudinal axis and wheel rotation speed varying are given in both time and frequency domains. All dynamic simulation results are similar to the results in reference [13] in which effective road plane inputs are used. Because of lacking of experimental results to compare with, the\ncomparison between simulation and experimental results of the same tire rolling over perpendicular cleat of 0.02m height with 2km/h speed is given. It can be seen that the amplitudes of calculated vertical and longitudinal response forces agree with the experiments very well. But there is a problem exists in time process, which is to be analyzed and very important for test validation of simulation results. In the end, an example of simulation of tire rolling over a series of continuous cleats is performed. But our results still have to be verified by experiments.\nCAMBER PROPERTY TIRE MODEL\n2008-01-1408\nThe 3-Dimensional Modal Parameter Tire Model and Simulation of Tire Rolling Over Oblique Cleats\nChengjian Fan, Dihua Guan and Baojiang Li Tsinghua University", + "Where: fr - radial force vector; ft - tangential force vector; Hrr, Hrt - radial and tangential displacement mobility matrix by radial excitation; Htr, Htt - radial and tangential displacement mobility matrix by tangential excitation.\nAccording to the modal theory, if l is the excitation point and p is the response point, transfer function with N orders is:\n2 1 j\nN lk pk\nlp k k k k\nH K M C\n(4)\nWhere, lr and pr , rK , rm , rC are mode shape, modal stiffness, modal mass and modal damping factor respectively.\nThe tread radial hr and tangential ht deformation in the contact patch are:\nr n r\nt m t\nh f k\nh f k\n(5)\nWhere: fn- vertical force; fm- longitudinal force; kr - normal stiffness of tread; kt - shear stiffness of tread.\nIn camber model, in which lateral modal parameters of tire are taken in to account, formula (3) becomes formula (6), and the total radial r, tangential t and lateral l deformation displacements are:", + "The case of simulation is referred to the example in reference [13]. The simulated tire is 205/55R16. Figure 4 shows some simulation results. The simulated operating conditions are: an oblique cleat (10*50mm2 strip) mounted on drum surface at an angle of 43\u00b0; drum diameter 2.5m; tire inflation pressure 250kpa; fixed axle height under vertical load 4000N; running speed 60km/h. It can be seen that the simulation results are similar to the ones in reference [13].\n0.3 0.4 0.5 -1000 -800 -600 -400 -200\n0 200 400 600 800\n1000\nFx [N\n]\ntime [s]\n0 20 40 60 80 100 120 10-6 10-5 10-4 10-3 10-2 10-1 100 101 102\nfrequency (Hz)\nS Fx\n0.3 0.4 0.5\n-200\n-100\n0\n100\n200\n300\nFy [N\n]\ntime [s]\n0 20 40 60 80 100 120 10-8 10-7 10-6 10-5 10-4 10-3 10-2 10-1 100 101\nfrequency [Hz]\nS Fy\n0.3 0.4 0.5\n-500\n0\n500\n1000\n1500\n2000\n2500\nFz\n[N ]\ntime [s]\n20 40 60 80 100 120 10-6 10-5 10-4 10-3 10-2 10-1 100 101 102 103 104\nfrequency [Hz]\nS\nFz\n0.3 0.4 0.5 -0.8 -0.6 -0.4 -0.2 0.0 0.2 0.4 0.6 0.8 1.0 1.2 [r ad /s ]\ntime [s]\n0 20 40 60 80 100 120 10-11\n10-10\n10-9\n10-8\n10-7\n10-6\n10-5\nfrequency [Hz]\nS" + ] + }, + { + "image_filename": "designv11_92_0003917_978-3-030-12346-8_35-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003917_978-3-030-12346-8_35-Figure4-1.png", + "caption": "Fig. 4. New patterns. Connecting rod (a), piston (b), anvil (c) and crankshaft (d)", + "texts": [ + " 2b) and the bell with a massive crack (Fig. 2c). So, the new patterns need to be didactic, cover different types and represent parts that are manufactured with sand casting process in the current industry. At the same time, there are some restrictions to take into account, like volume (V) and size. These limits are established by the laboratory material, more specifically the furnace (V = 1 dm3) and the flask (150 mm long, 200 mm wide and 100 mm high), respectively (Fig. 3). The first pattern considered is a scaled connecting rod (Fig. 4a). It replaces the anchor since they have similar characteristics: a flat piece with 2 holes in the demolding direction. The second pattern considered is a piston. This full pattern replaces the bell. The modifications in this case are the elimination of the compression rings and bolt holes, generally located in the flat side areas (Fig. 4b). The casting part will be subjected to other forming processes (machining) to achieve the final shape. The third chosen pattern is an anvil and will obviously replace the previous one. The new shares its design characteristics with its predecessor (Fig. 4c). The last pattern selected is a crankshaft. It stands out for a more intricate design, with abrupt changes of section which may result in several defects (shrinkage cavities, lack of full mold filling, etc.) if the gating system is not correctly design (Fig. 4d). Many of the defects obtained in casting processes are originated due to a bad gating system design. For its design and calculation, there are diverse analytical methods that can be applied to obtain an optimal sand casting process. Assuring a complete and correct mold filling is essential to avoid part defects. This can be achieved with a proper calculation of the feeding channels and risers, among others [9]. For the design of all the patterns, the SolidWorks 2016 software has been used. It can be especially highlighting the tool \u201cExit Angle Analysis\u201d, a graphical tool that allows to detect easily the pattern faces that needs an exit angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002877_s0027131418050127-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002877_s0027131418050127-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of FEPS.", + "texts": [ + "4) from Aspergillus niger (Sigma, Germany) was used in the form of a lyophilized protein with a declared activity of 136.3 U/mg; glucose (Merck, Germany) was used for the preparation of model solutions. Perfluorosulfonated polymer (PFS) (10% solution in isopropyl alcohol, Plastpolimer, Russia), a structural analogue of Nafion, was used for immobilizing GOx. The activity of the enzymes was controlled spectrophotometrically according to the techniques described in [17]. Four-electrode screen-printed structures (FEPS) manufactured in-house using the instrumentation SCF 550 (China) (Fig. 1) were used as a base for the production of a multibiosensor. The electrodes were printed on a polyethylene terephthalate film (Vladimirskii Khimicheskii Zavod, Russia) using a silvercontaining polymer paste (NPP Delta-Pasty, Russia), carbon paste (Gwent Electronic Materials, United Kingdom), and a UV-curable insulation paste (UNICA, Belgium). Modification of the working electrode surface with PB was carried out according to [18]. The GOx was immobilized according to [11] and the LOx was immobilized according to the technique in [13]", + " The analyte concentration was determined from the graduation graph constructed using the results of the measurements of the model solutions as described above. In order to avoid the possible sensitivity loss of the biosensor as a result of contact with the sample, a control measurement of the model solutions of analytes with concentrations of 10 and 50 \u03bcM was carried out after the analysis of each sample. The obtained data were compared with the data presented in the certificate of the standardized sample, as well as with the data obtained on an EcoBasic commercial glucose and lactate analyzer (Care Diagnostica, Germany). A FEPS (Fig. 1) was developed during the construction of a multibiosensor for the simultaneous determination of glucose and lactate. It comprised two working electrodes with their surface modified with sensor materials ensuring the selective determination of the respective analyte. Two successful previously developed techniques for the modification of the surface of the working electrodes with PB and immobilization of the GOx and LOx from the solutions with a high content of the organic solvent into Nafion and siloxane membranes, respectively, were used during the construction of the sensor material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001940_aset.2018.8379848-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001940_aset.2018.8379848-Figure1-1.png", + "caption": "Fig. 1. Quadrotor UAV coordinate system", + "texts": [ + "ndex Terms\u2014Quadrotor UAV, Modelling, Internal Model Control IMC, Disturbance, Uncertainty. I. INTRODUCTION Quadrotor UAV is classified with VTOL aircraft category, it is used in many fields like commercial delivery, rescue, scientific research, photography due to its performances and capability of hover and high payload. The quadrotor contains two perpendicular arms on 4 propellers holding a motorpropeller on its end as shown in Fig. 1. The motion of the quadrotor can be controlled through varying motors speed. In fact changing the speed of the motor 1 and 3 produces the rotation around Y axis (Pitch). To rotate around X axis we can adjust the speed of motor (2) and (4). To rotate the quadrotor around the Z axis (Yaw), we can increasing or decreasing motors (2) and (4) versus motors (1) and (3). Changing simultaneously the motors speed results a vertical motion. [1] [2] Modeling of the qudrotor UAV is developed with three principal approach, [3] used Newton-Euler and Lagrange formalism whereas [4] used quaternion method", + " Section III presents the IMC control strategy and the calculation of IMC Controller to stabilize the quadrotor position and attitude. Simulation results and discussion is shown in Section IV. Finally in Section V future work and conclusion are given. This section develop the mathematical model of the quadrotor using Newton-Euler formalism. The generalized coordinates of the quadrotor are defined as: q = (x, y, z, \u03c6, \u03b8, \u03c8) \u2208 6 (1) An illustration of the generalized coordinates of the quadrotor is shown in Fig 1. where E fixed inertial frame, B the quadrotor body frame. Denoting \u03be = (x, y, z) is the position vector of the quadrotor center of mass relative to E, the quadrotor angles are given by \u03d1 = (\u03c6, \u03b8, \u03c8), where \u03c6 is the roll angle around the x-axis, \u03b8 is the pitch angle around the y-axis and \u03c8 is the yaw angle around the z-axis. Thus the transformation from the fixed frame E to the body frame b is given the following matrix R = \u23a1 \u23a3c\u03c8 c\u03b8 c\u03c8 s\u03b8 s\u03c6\u2212 s\u03c8 c\u03c6 c\u03c8 s\u03b8 c\u03c6+ s\u03c8 s\u03c6 s\u03c8 c\u03b8 s\u03c8 s\u03b8 s\u03c6+ c\u03c8 c\u03c6 s\u03c8 s\u03b8 c\u03c6\u2212 s\u03c6 c\u03c8 \u2212s\u03b8 c\u03b8 s\u03c6 c\u03b8 c\u03c6 \u23a4 \u23a6 (2) Noting that s\u03b8 = sin(\u03b8) and c\u03b8 = cos(\u03b8)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003662_978-3-030-13273-6_43-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003662_978-3-030-13273-6_43-Figure8-1.png", + "caption": "Fig. 8. Point cloud visualisation of the reference measurement.", + "texts": [ + " The combination of two scanning sessions also requires selecting a minimum of 3 recognised markers common to both. Selection of a larger number of markers results in averaging the relative position of the point clouds with each other by minimising the mean square error, which usually leads to an increase in the accuracy of mutual positioning. As a result of the above operations, a dense cloud consisting of several million points was created and initially presented as a surface consisting of triangles stretched between the three points nearest to each other. The result is shown in the figure below (Fig. 8). Nevertheless, these data are not suitable for further processing and require another operation \u2013 polygonisation, which means further processing with a number of tools for filtration of the point cloud and generation of a surface. The data acquisition process generates a large volume of data points, which is a big problem in further processing due to the limited computing capabilities of computers. To be able to perform further reconstruction procedures, the point cloud should be converted into a set of surfaces, without reducing the accuracy of the mapping" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000316_mawe.200800287-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000316_mawe.200800287-Figure12-1.png", + "caption": "Figure 12. Schematic showing potential arrangement of double sided forming of SPF component using a laser forming SPF cell. (Courtesy LISTechnology Ltd)", + "texts": [ + " Incidentally, in a recent Simulation Study we proved: Part Cost reduction 60 %, Productivity increase 260 % (over twice the number of parts produced at less than half the cost) 4. Green technology - Energy Efficient - Environmentally Friendly - Low Cost - Waste Free Activity - High Value Added = Safe, Sustainable Manufacturing The team at the University of West of England led by Professor Alan Jocelyn are very hopeful that they will be able to raise the capital required to build the prototype machinery (Fig. 12) that will enable them to prove the concept by applying the technology to the manufacture of components required in a range of sectors. Mat.-wiss. u. Werkstofftech. 2008, 39, No. 4-5 Overview \u2013 Superplasticity Community 273 Nowadays, a Roadmap seems to be associated with every major technology, pointing the way forward and highlighting areas of growth and trends in the market. Superplasticity should be no exception, and members of the SPF Community in Europe have begun turning their attention to a Roadmap for superplasticity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002221_978-3-319-99270-9_27-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002221_978-3-319-99270-9_27-Figure1-1.png", + "caption": "Fig. 1. Rotor under investigation", + "texts": [ + " In this study, the recorded behavior of the studied rotor in sudden failure of the electromagnetic field is demonstrated. The description of the support stiffness and friction coefficients in the bearings are fine-tuned in the rotor model to meet the measured results during the dropdown. In addition, the fine-tuned rotor-system model is used for studying the contact force and contact stress in the backup bearing. Furthermore, the rotor orbit and the displacement of the rotor during the dropdown are featured. Figure 1 shows the rotor under investigation. The rotor is modeled using shear deformable beam elements. In a rotor dropdown, the backup bearing supports the rotor by nonlinear contact load. The equation of motion is as follows [12]: M\u20acX\u00fe C\u00fexG\u00f0 \u00de _X\u00feKX \u00bc x2F1 \u00feF2 \u00f01\u00de where M, C, K, and G are mass, damping, stiffness and gyroscopic matrix, respectively. X is the vector of the generalized coordinate. The vector F1 is the vector of nodal unbalance and F2 represents the vector externally applied forces. The angular velocity of the rotor is x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000118_20080706-5-kr-1001.00015-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000118_20080706-5-kr-1001.00015-Figure1-1.png", + "caption": "Fig. 1. Relative position of the reference path and the plant", + "texts": [ + " At the last, the effectiveness of the method are shown by numerical examples with an automobile model. It is assumed that plants move in horizontal plane and a reference path is given in the plane. Suppose the reference 978-3-902661-00-5/08/$20.00 \u00a9 2008 IFAC 90 10.3182/20080706-5-KR-1001.1833 path is given as a function with respect to the path length s, i.e., the reference path is represented as Pr(s) = (\u03ber(s), \u03b7r(s)) T . In particular, we also assume \u03ber(s) and \u03b7r(s) belong to class C3. \u03a3 is inertial coordinate and orthonormal bases of \u03a3 is defined as e1, e2(See Fig. 1). Pr(s) is written as follow: Pr(s) = \u03ber(s)e1 + \u03b7r(s)e2 (1) Then, the curvature \u03bar of each point s in reference path is represented by \u03bar(s) = d\u03b8r(s) ds . (2) Since \u03ber(s) and \u03b7r(s) belongs to class C3, \u03bar is differentiable in each s. We consider path following control for a plant which is given by the following non-linear state equation: x\u0307p(t) = fp1(xp(t)) + fp2(xp(t))u(t), (3) where xp(t) \u2208 Rn is state and u(t) = [ u1(t), \u00b7 \u00b7 \u00b7 , um(t) ]T\u2208 Rm is control input. { \u00b7 } is the differential operator with respect to t" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001724_icit.2018.8352228-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001724_icit.2018.8352228-Figure2-1.png", + "caption": "Fig. 2 Cross\u2013sectional view of consequent-pole-type motor with dynamic eccentricity due to the unbalanced vibration. In addition, the d-axis current is injected to change the UMF.", + "texts": [ + " In the elastic inboard rotor, an unbalanced force acts on the rotor and the unbalanced vibration is the dominant vibration during the rotation. This dynamic eccentricity induces the UMF on the rotor. This paper proposes to use AC current injection into the daxis to generate the radial excitation force in a PM motor. The d-axis current can change the amount of the bias magnetic flux generated by the PMs. The fluctuation of the amount of the bias magnetic flux results in the radial excitation force. Fig. 2 shows the cross sectional view and the flux flow in a PM motor with the eccentricity in the positive X direction. A consequent-pole-type motor with four PMs in the rotor was chosen for this study. The PMs in the rotor generate the bias magnetic flux. In the clearance of the positive X direction, the magnetic flux is strengthened by the eccentricity as follows: = +\u2212 2 (1) where is the magnetic permeability of free space, is the magneto-motive force of the PM, is the number of turns in the d-axis windings, is the d-axis current, and is the gap including nominal air gaps and the PM thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003470_amcon.2018.8614963-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003470_amcon.2018.8614963-Figure2-1.png", + "caption": "Fig. 2 Five-axis machine tool configuration", + "texts": [ + " Production Scheduling Processing efficiency optimizationProcess simulation thermal compensation Machine fault detection and diagnosis system SaaS CNC work piece weight parameter matching Machine status detection Tool management system Cloud computing NC program management system Conversational NC program editing Tool measurement and correction Virtual measurement Five-axis rotation center error measurement and compensation Processing origin setting Processing mode (accuracy, quality, efficiency) EI-PaaS Device Management ( Condor ) Metering DB Services AFS Management Center Knowledge Center Security DevOps SSO Prognostic & Health Management Defect Pattern Analysis OpenStack Visualization service Microsoft Azure Public Cloud Module PaaS Private cloud module PaaS Edge computing Edge NC program management system Conversational NC program editing Feed rate adaptive control Chatter avoidance Thermal compensation Spindle monitoring Machine measurement and compensation Virtual measurement Production Scheduling Process simulation Tool management system Processing mode (accuracy, quality, efficiency) Five-axis rotation center error measurement and compensation NC program interference check (prevention of collision) Tool measurement and correction Processing efficiency optimization Machine fault detection and diagnosis system CNC work piece weight parameter matching Machine status detection Tool life management and damage detection Processing origin setting Machine measurement and compensation Fig.2 Machine tools application service module requirements data analysis & visualization Platform operation & maintenance Device connectivity & data management Inform ation security operations support Data Analysis Automated testing Paid service Data visualizationComplex Event Engine App and service management Data management & storage Equipment management SCM (supply chain management) CRM (Customer relationship management) ERP (Enterprise resource planning) MES ( Manufacturing Execution System) Mechanical End Controller Layer Acceleration gauge, temperature, current PLC Industrial Computer Switching board CNC Controller Production management digitization (from industry 2", + " [5] Makoto Fujishima, Katsuhiko Ohno, Shizuo Nishikawa, Kimiyuki Nishimura, Masataka Sakamoto, Kengo Kawai, Study Of Sensing Technologies For Machine Tools, Cirp Journal Of Manufacturing Science And Technology 14 (2016) 71\u201375. [6] Jay Lee, Edzel Lapira, Behrad Bagheri, Hung-An Kao, Recent Advances and Trends In Predictive Manufacturing Systems In Big Data Environment, Manufacturing Letters 1 (2013) 38\u201341. [7] Jay Lee, Behrad Bagheri, Chao Jin, Introduction to Cyber Manufacturing, Manufacturing Letters 8 (2016) 11\u201315. 120 ISBN: 978-1-5386-5609-9 Figure 2 shows the investigated five-axis machine tool with two rotary movements about the C-axis of the table and B-axis of the spindle. Tool axis orientation is necessary for postprocessor development. The data represented in Table 1 does not contain the tool axis orientation which should be determined further. Generally, tools are divided into ball mill, end mill and bull nose mill. This paper would use the ball mill as an example to explain the mathematical module. The five-axis tool path file provided by Mastercam will include the tool tip vector 1 T x y zQ Q Q Q , the contact point vector T x y z 0N N N N , and the contact point vector 1 T x y zP P P P " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000779_icit.2009.4939600-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000779_icit.2009.4939600-Figure4-1.png", + "caption": "Fig. 4. The full 3-D model of the flameproof motor", + "texts": [ + " The resulting differential equations are: [ ] { } { } [ ] [ ] { } [ ] { } { } \u03a9=\u00d7\u2207\u00d7\u2212\u2207+ \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 \u2202 \u2202+\u22c5\u2207\u2207\u2212\u00d7\u2207\u00d7\u2207 inAVV t AAvAv e 0\u03c3\u03c3 \u03c3 (2) [ ] [ ] { } [ ] { } { } \u03a9 =\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u00d7\u2207\u00d7+\u2207\u2212 \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 \u2202 \u2202\u22c5\u2207 in AVV t A 0\u03c3\u03c3\u03c3 (3) [ ] ( ) ( ) ( ){ }3,32,21,1 3 1 3 1 vvvvtrve ++== (4) where V is electric scale potential, \u03c3 is conductivity, \u03bd is reluctivity, \u03a9 is conducting region. Since in the end region, the motor has 3-D magnetic flux. It is necessary to use 3-D finite element analysis (FEA) for accurate determination of eddy current field. The model for field solution in the end region is shown in Fig. 4. It has two parts, circuit part and finite element part. The circuit part is used as magnetic field excitation and the finite element part is used for field analysis. Fig. 5 shows the no load magnetic distribution in the end region. C. 3D Solid Source Conductor Inductance Calculation The typical structure of the solid source conductor is illustrated in Fig. 6. The magnetic flux density B is a vector. It can be expressed as in (5) kzyxRjzyxQizyxPzyxB ),,(),,(),,(),,( ++= (5) where P, Q and R have consecutive first order partial derivative" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002493_s11277-018-5949-1-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002493_s11277-018-5949-1-Figure3-1.png", + "caption": "Fig. 3 When landmarks and are observed, the regional coordinate system is created by landmark", + "texts": [ + " If the recognized landmark did not be in a previous memory, a new node is established and a regional coordinate system is set up. If it is node , the regional coordinate system 1 3 is represented by \u03a9 = \u27e8 v, \u27e9 . The representation of the regional coordinate system creates the spatial landmark relations as follows. 4. In this case, the robot makes spatial landmark relations about the recognized landmarks around it. If no other landmarks are observed, then the semantic mapping is complete with the current node. Figure\u00a03 shows the results, which can be expressed as 5. As indicated by (d) in Fig.\u00a04, the robot approaches to the new landmark to estimate distance and angle. If the landmark did not find on a previous memory, a new node is established, as showed by the regional coordinate system for (e) in Fig.\u00a05. 6. The robot creates spatial landmark relations of the landmarks at position (e) in Fig.\u00a04. If no more landmarks appear, the robot has completed its semantic-metric map with the current node. { \u0398 , S } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v )}} 1\u0393 = { \u0398 , S , } a,b\u22081O{ \u0398 , S , } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v ) , ( 1 , 1 )}} { \u0398 , S , } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v ) , ( 1 , 1 )}} 2\u03a9 = \u27e82v, \u27e9 \ufffd \u0398 , S \ufffd = \ufffd \ufffd 2 , 2v \ufffd , \ufffd \ufffd 2 , 2v \ufffd , \ufffd 2 , 2v \ufffd\ufffd\ufffd 1 3 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000880_iciea.2008.4582672-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000880_iciea.2008.4582672-Figure3-1.png", + "caption": "Fig. 3 The fuzzy control block", + "texts": [ + ", the uncertainties both structured and unstructured are all ignored, then (9) reduces to a decoupled linear time-invariant system: \u03b8 u= (15) Equation (15) shows that perfect dynamic decoupling and global linearization of the HPMSM can be actualized if the model errors and external perturbations are eliminated. Then we propose a new algorithm by introducing fuzzy control into the CTM structure to resolve the uncertainties problems. In the proposed algorithm, the PD control part +p dK e K e , which is part of servo portion u, is taken place by 3 fuzzy control blocks, each of which contains a 2- dimensional fuzzy controller [11]. Fig. 2 shows the control diagram of the algorithm proposed in this paper, and Fig. 3 shows the structure of the fuzzy control block in the diagram. Pg 1032 Each fuzzy controller is responsible for the control of one position axis. Both ek and eck are quantitative factors, which converse e and e into the input universe of the fuzzy controller. uk is the proportional factor, which translates the output universe of the fuzzy controller into actual output. The input universe and the output one are both set as [ 6,6]\u2212 . Seven language sets, each of which is corresponding to a membership function, are defined in the universe as {PB, PM, PS, Z, NS, NM, NB}" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002434_s1068798x18080087-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002434_s1068798x18080087-Figure3-1.png", + "caption": "Fig. 3. Supporting roller of a four-high stand in a highspeed rolling mill [3].", + "texts": [ + "eywords: radial clearance, contact angle, contact deformation, radial and axial dynamic loads DOI: 10.3103/S1068798X18080087 A radial ball bearing of 0000 type may be used to withstand purely axial loads. As an illustration, we show the input shaft of a cylindrical gear in Fig. 1 (from [1]); the input shaft of a bevel gear in Fig. 2 (from [2]); and the supporting roller of a four-high stand in a high-speed rolling mill in Fig. 3 (from [3]). Each shaft is mounted on two radial roller bearings intended to withstand the radial load and one radial ball bearing absorbing the purely axial load. This is ensured by a hole drilled in the housing under the bearing, whose diameter exceeds that of the bearing\u2019s external race. The radial ball bearing is used in horizontal and vertical electrical machines, even with relatively large axial loads on the shaft created by the drive. Radial thrust bearings of type 6000, 36000, 46000, and 66000 are not used in that case since the radial ball bearing operates more rapidly and with less noise, thanks to the fewer shape errors in its outer raceway and to the incorporation of a separator such that all of the external surface is centered with respect to the bearing\u2019s outer race" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000156_msf.618-619.387-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000156_msf.618-619.387-Figure2-1.png", + "caption": "Figure 2. The mould design", + "texts": [ + " The casting grade of aluminium, A356 is selected as the experimental material due to the ease of casting and ready availability. Tensile strength and surface roughness being the primary responses, the moulds are essentially designed to cast the tensile test specimens 388 Light Metals Technology 2009 of standard shape and size as per the British Standard BS 10:002 1990. The moulds are designed using Solid Works and a couple of samples are included in each mould, connected to a carefully designed runner, filter, gating and riser system as shown in Figure. 2, in order to reduce printing expenses. Considering the practical difficulties in clearing the powder from internal cavities, the moulds were finally designed to be printed in two halves, with horizontal parting surfaces, as shown in Figure. 3. The metal to be poured is heated to around 7000C, as per the recommendations by the local foundry and manually filled into the printed moulds. The specimens are then cut from the casting and analysed for the surface quality and tensile strength. Figure 3. Printed moulds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000182_09544062jmes1161-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000182_09544062jmes1161-Figure1-1.png", + "caption": "Fig. 1 Bearing pedestal model with eight possible coefficients", + "texts": [ + " The SCM described here is quite similar to the influence coefficient method developed by Nikolajsen and Holmes [21]. However, instead of using the flexibility matrix as in the influence coefficient method, the stiffness of the shaft is adopted. The objective of these investigations is to validate the estimated bearing pedestal coefficients, as well as the SCM program using the commercial ANSYS Inc. v11.0 software as a tool to compare the results. 2 ANALYTICAL FORMULATION 2.1 Parameter estimation of bearing pedestal Referring to Fig. 1, consider a bearing pedestal of mass m supported by the direct and cross-axis springs and dampers modelled by eight linearized coefficients, a perturbation force f is applied at a rotated angle \u03b1 from the horizontal direction. The displacement, velocity, and acceleration vector quantities for the equations of motion in the x- and y-directions are mx\u0308 + Cxxx\u0307 + Cxy y\u0307 + Kxxx + Kxy y = \u03b2xf (1) my\u0308 + Cyy y\u0307 + Cyxx\u0307 + Kyy y + Kyxx = \u03b2y f (2) Equations (1) and (2) can be expressed in the form of frequency domain (\u2212m\u03c92 + Cxxj\u03c9 + Kxx)X (j\u03c9) + (Cxy j\u03c9 + Kxy)Y (j\u03c9) = \u03b2xF (j\u03c9) (3) (\u2212m\u03c92 + Cyy j\u03c9 + Kyy)Y (j\u03c9) + (Cyxj\u03c9 + Kyx)X (j\u03c9) = \u03b2y F (j\u03c9) (4) where f is the multi-frequency force signal applied to the rotor, \u03b2x = b cos \u03b1, \u03b2y = b sin \u03b1, \u03b1 = tan\u22121(\u03b2y /\u03b2x) is the forcing angle, and b is the input force transducer\u2019s gain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003418_ieem.2018.8607695-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003418_ieem.2018.8607695-Figure7-1.png", + "caption": "Figure 7: -&# # $, '", + "texts": [ + " In this case the beam has traveled from X=0 to X=4 mm and the picture is taken when the beam is at X=4mm. In this figure limited difference between FE considered as the reference and abacus that needs to prove its reliability can be observed. In 3 different points, a mean difference of 58\u00b0C, a difference between the 2 curves of 4.48% and a very close thermal history have been noticed. For the second case study the goal is to determine the temperature reached when melting a prismatic part as represented fig. 7. The trajectory used is a simple hatching as represented fig. 8 and the beam parameters and travel speed are constant as given table 1. Because mmeasuring the temperature in the built chamber has not yet given satisfactory results, the simulation results have been compared with a real built. For that, a part has been built with the proposed strategy and part defects and maximum reached simulated temperature has been compared. The optimal temperature range to melt the Ti6Al4V powder has been determined to be between 2700\u00b0C and 2760\u00b0C by [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002159_1.5046268-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002159_1.5046268-Figure3-1.png", + "caption": "FIGURE 3. Product Geometry of (a) ASTM D638-03 Type V and (b) ASTM D256-04", + "texts": [ + " The barrel temperature is bound by low temperature that result in insufficient melting material and high temperature that leads to over-melting and material degradation. The injection pressure is bound by low pressure that leads to short shot or unfilled cavity problem and high pressure that result in flash [5]. According to the molding window result as shown in Figure 2, the parameter settings which are chosen for injection molding process is shown in Table 2. Products made in this injection molding process are test specimens of tensile and impact test with ASTM D63803 Type V and ASTM D256-04 standard as shown in Figure 3. Injection molding machine used in this experiment is HAITIAN-MA900/260e. Furthermore, the products will be tested such as tensile test and impact test, also the crosssection of the specimen was observed by using SEM. Several types of testing were performed on injection molding products, such as tensile test, impact test and scanning electron microscope (SEM) investigation. Mechanical testing in the form of impact test and tensile test are conducted to know the characteristic of mechanical properties of bio-composite material" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001718_icit.2018.8352487-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001718_icit.2018.8352487-Figure1-1.png", + "caption": "Fig. 1. Concept model of proposed robot", + "texts": [ + " It becomes substitute of pitch control function in rotor mechanism. By this manner, it is possible to control the spatial moving in forward direction. By this concept, this robot can expand its working space and expand working time because this robot basically moves by using a wheel which has a high energy efficiency for moving. In particular, this paper introduces the prototype, discussed the study on energy efficiency and experimental results of the estimation of direction in x axis (forward direction). The prototype [5] of the proposed system is shown in Fig. 1. This system is composed of a pair of rotors and motors, single in-wheeled DC motor, three ball rollers supporting the body. The wheel is attached to the body by aligning its grounding point with the rotation center of rotors in order to change the body orientation by counter torque of rotors. Because single wheel system is not stable, three auxiliary legs equipped with ball roller at their tip are adopted for supporting the body. The proposed mobile system uses coaxial rotors which enables spatial moving and can be implemented in compact space", + " The horizontal length of this stair is 1.95 [m]. The initial velocity is set to 0.5[m/s] and flying horizontal distance is set 2.0 [m], and then the flying time is calculated as follows by using uniform motion model: t = x v0 = 2.0 0.5 = 4.0[s] (15) The robot is used in this experiment is shown in Fig. 5. The robot is combined of Parrot AR.Drone 2.0 and 3 wheels robot. Each mobile mechanism take charge of spatial moving and moving on ground. This robot is bigger than the proposed prototype shown in Fig. 1. Therefore the coefficient of air friction and projected area of Fig. 5 is bigger than the prototype (Fig. 1). Thus if the position damping can be neglected using AR Drone, the position damping can also be neglected in various size of mobile robot. Therefore the robot shown in Fig. 5 is adopted as benchmark of worst case of air friction. The distance of the robot is measured by kinect for windows v2. The runway(0.5[m]) is prepared for maintaining initial velocity 0.5[m/s] at take off. When the robot flies, the rotors never generate thrust force for moving forward, the robot moves by using only inertial force made by wheel motor on the runway" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure8-1.png", + "caption": "Fig. 8. Third metamorphic limb with phase e1", + "texts": [ + " The metamorphic limb is synthesized with the phase e2 in Fig. 7. Link a is fixed to base, and rotational joint of axis 1 is locked. Axis 2 and axis 3 are perpendicular each other which is equivalent to Hooke joint with u and v rotate axis direction. Axis 4 and axis 5 are parallel to axis 6 with w3 direction. Then metamorphic limb takes on mobility five which can realize motion type of submanifold T2(w3) \u00b7S(p3). Keeping axis 1 is locked and rotating link c so that axis 2 and axis 3 with u and w rotate axis direction in Fig. 8. Axis 4 and axis 5 are parallel to axis 6 with v3 direction. Then metamorphic limb satisfy form of R(q31,u) \u00b7R(q31,w) and R(q32,v3) \u00b7R(q33,v3) \u00b7R(q34,v3) which generated by T2(v3) \u00b7 S(q3) that have mobility five and can realize motion type of such submanifold. In this section, metamorphic mechanisms will be assembled to satisfy motion type of 1R2T and 2R1T. Motion type of 1R2T with the mobility three, which is intersection of motion of limbs. In other words, constraint of moving platform is union of constraint of limbs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000752_sice.2008.4655087-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000752_sice.2008.4655087-Figure8-1.png", + "caption": "Fig. 8 Definition of \u03c8off .", + "texts": [ + " The double legs stance periods beginning at the contacts of the slow leg and the fast leg are named as \u2018slow\u2019 and \u2018fast\u2019 DLSP, respectively. There is adaptive control results from the cerebellum to enable human to walk on a splitbelt treadmill. Adaptive control using contact sensors was applied to enable Tetsuro to walk on a splitbelt treadmill. Hosoda et al.[9] and Tsuchiya et al.[10] achieved adaptive walking by adjusting the solenoid valves of pneumatic motors based on contact information and by feeding back contact information to neural oscillators. A schematic view of one step is shown in Fig. 8. In Fig. 8, \u03c8off is the hip joint angle of the swinging leg when leaving the ground. Getting off the ground occurs during the single leg supporting phase. The p-gain of the hip joint of the supporting leg is updated for the nth step in the adaptation and post-adaptation period using the following equation. k sp h p(n) = kag\u00d7(\u03c8\u0304off\u2212\u03c8off (n\u22121))+k sp h p(n\u22121) (4) kag is the gain of P-gain adjustment. \u03c8\u0304off is the hip joint average angle of the swinging leg when leaving the ground in the baseline period. k sp h p(n\u2212 1) is the hip joint p-gain of the supporting leg on the n \u2212 1th step. Putting this adaptive control, Tetsuro can adjust the stiffness of the supporting leg according to the change of the treadmill speed. We made Tetsuro walk in the experimental setup described in Section 5.1 . Tetsuro autonomously detected the switch from the pre-adaptation stage to the adaptation stage by the change of lift off timing (Fig. 8), and started P-gain adjustment expressed by eq.(4). Tetsuro did not change its control method at the switch from the adaptation stage to the post-adaptation stage. Two indexes, the stride length and the duty ratio, measured in a robot experiment are shown in Fig. 9. The same indexes measured in human (normal subject) experiments are shown in Fig. 10 (modified from Morton et al.[4]). Each circle in those figures means the value of the index measured at the walking cycle. In both the stride length and the duty ratio, data of a robot experiment and data of human experiments show similar pat- - 2511 - terns and similar values except for the stride length because of the difference of the leg length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003161_978-3-030-03047-6_8-Figure8.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003161_978-3-030-03047-6_8-Figure8.4-1.png", + "caption": "Fig. 8.4 Physical model of the active suspension with seated human body (a) and experimental test stand (b)", + "texts": [ + "12) with the following matrices: C2i = [ 1 0 0 0 \u00b7 \u00b7 \u00b7 0 0 0 1 0 0 \u00b7 \u00b7 \u00b7 0 0 ] (8.13) D2si = [\u22121 0 0 0 ] D2ai = [ 0 0 ] (8.14) A real-time control problem is formulated using the measurement vector yi(t) and multiplying it by the output feedback gain vectorKi = [k1i, k2i] in the following order: Fai(t) = Kiyi(t), i = x, y, z (8.15) where k1i is the proportionality factor of the relative displacement feedback loop, and k2i is the proportionality factor of the absolute velocity feedback loop. In Fig. 8.4, physical and mathematical model of the horizontal seat suspension with seated human body is shown. An elementary 3-DOF model is employed to represent the system bio-dynamic response under harmful vibrations [4]. In order to evaluate the desired force that should be generated in the system actively, the linear spring csx and damper dsx are utilised to describe essential characteristics of the horizontal seat suspension. The linear suspension system can be modelled using a state-space approach for the purpose of the primary controller synthesis. The following state variables are used to reproduce the movement of a human body in the longitudinal x-direction: xx(t) := [q1x, q\u03071x, q2x, q\u03072x, q3x, q\u03073x]T (8.16) where q1x, q2x, q3x and q\u03071x, q\u03072s, q\u03073x are the displacements and velocities of the biomechanical model (Fig. 8.4). The external disturbances are defined as the displacement qsx and velocity q\u0307sx of input vibration: wsx(t) := [qsx, q\u0307sx]T (8.17) The state-space equation of the active horizontal suspension is obtained using the applied force Fax as an output from the primary controller: x\u0307x(t) = Axxx(t) + Bsxwsx(t) + BaxFax(t) (8.18) The state (system) matrix is formulated in the following form: Ax = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1 0 0 0 0 \u2212 c11x m1 \u2212 d11x m1 c12x m1 d12x m1 0 0 0 0 0 1 0 0 c12x m2 d12x m2 \u2212 c22x m2 \u2212 d22x m2 c2ni m2 d2ni m2 0 0 0 0 0 1 0 0 c23x m3 d23x m3 \u2212 c33x m3 \u2212 d33x m3 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003726_iwed.2019.8664372-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003726_iwed.2019.8664372-Figure1-1.png", + "caption": "Fig. 1. Navigation frame and body frame, with the motor positions and direction of rotation.", + "texts": [ + " In this work a multirotor aircraft with no tilted motors is referred to as a \u201cnormal configuration\u201d, where the 978-1-5386-9453-4/19/$31.00 \u00a92019 IEEE rotational axis of each motor is parallel to Zb axis. In the \u201cdihedral configuration\u201d, all motors are equally tilted around Zb and Yb axis towards the centre of gravity (14 deg.), other than in the V-tail configuration, where the rear motors (M2 and M4) are tilted around Xb axis towards each other (30 deg.). To model a multirotor aircraft it is seen as a rigid body, moving in the coordination of the navigation frame with the index n (Fig. 1). Due to the small distances travelled by the multirotor aircraft, earth\u2019s curvature is neglected [11]. The coordination system of the object is called body frame with the index b, in this case the Quadcopter (Fig. 1). The conversion of different force vectors between navigation frame and body frame, Euler angles and rotation matrices are used as following. (1) where \u03d5 is Euler angle around X [rad]; \u03b8 is Euler angle around Y [rad]; \u03c8 is Euler angle around Z [rad]. Since the body frame is usually tilting in 3 directions at the same time a rotation matrix for all axes is needed using the rotation sequence roll \u03d5, pitch \u03b8, yaw \u03c8 [9], [13], [14]. (2) where is rotation matrix body to navigation frame (3). If the conversion needs to be done from the navigation frame to the body frame the transposed rotation matrix is used as following" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001637_j.proeng.2018.02.047-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001637_j.proeng.2018.02.047-Figure3-1.png", + "caption": "Figure 3 : contact pressure distribution due to interference", + "texts": [ + " The transmittable torque is extrapolated from the contact pressure at the interface with Coulomb\u2019s friction law, under the assumption that full sliding occurs (so that : shear stress at the interface = pressure*friction coefficient). The German literature (e.g. [3,11]) interestingly introduces a distinction between friction coefficient in axial and peripheral direction. 480 Samuel K\u0153chlin et al. / Procedia Engineering 213 (2018) 477\u2013487 4 K\u0153chlin et al./ Procedia Engineering 00 (2017) 000\u2013000 When the hub is not axisymmetric, the pressure distribution at the interface is obviously modified. In the case of our connection, the key slot in the hub causes a pressure drop in its vicinity (Fig.3). The corresponding sliding torque (calculated by FEA) is 13% lower than the value given by the analytical formula above (taking into account the reduced contact surface, but with a uniform pressure). Two additional effects are not considered in 2D models. The first one is the stress peak on the edge of the hub end, which is clearly visible in Fig.3 (top right red strip). This edge effect is also called \u201cpunching effect\u201d (since a shorter part is pressed against a longer one). The second one is the partial sliding occuring at the interface when enough torque is applied to the connection, even if it is still fully transmitted (so well below the full sliding torque) [6]. The combination of both is critical for fretting fatigue initiation. Several authors have investigated the effect of external loads on an interference fit. Ref.[10] includes a calculation procedure to evaluate the modification of pressure distribution in case of external load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003917_978-3-030-12346-8_35-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003917_978-3-030-12346-8_35-Figure6-1.png", + "caption": "Fig. 6. Gating system", + "texts": [ + " Figure 5 shows the process with which the exit angle for the crankshaft is provided. Due to the large number of faces perpendicular to the partition line, the software marks in yellow the ones that need an exit angle. All patterns, except the piston, have been divided in 2 parts (cope and drag) and a keyway is designed to make easier their attachment, following the usual sand casting procedure. To build the optimal part shape, an approximation to the different analytical methods of calculation for the sand casting elements has been made. The gating and feeding system (Fig. 6a) have been calculated with two different methods to ensure the correct filling of the mold (Table 1). The nomogram method [9], following the \u201cBritish Non-Ferrous Metals Technology Center\u201d guidelines, and the analytical method, implemented in EES software. A series of considerations must be taken in the design and calculation of the gating system. As the Fig. 6 shows, the system consists of a base for the sprue (Dsprue base), shaped as a cylinder of height (hsprue base) equal to twice the thickness of the casting channel (ec) and a diameter equal to twice the lower diameter of the sprue (D2) (Eqs. 1 and 2). In addition, the area of the runners must be twice the area of the pouring channel (Acc), the sprue has a conical shape and the casting channel will be square base, being its thickness the square root of the pouring channel section (Eq. 3). The maximum filling time has been calculated using the Chvorinov equation [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000627_isam.2009.5376920-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000627_isam.2009.5376920-Figure4-1.png", + "caption": "Fig. 4. Shape of the pin", + "texts": [], + "surrounding_texts": [ + "Two pins were made from tungsten carbide (WC-Co, AF310, Axismateria, Japan). An edge angled at either 30\u00b0 or 45\u00b0 was applied to the tip of the pins (see Figs. 3 and 4) to help with the insertion. The insertion force was measured for each pin. The pins were fabricated by grinding, and their diameters were measured with a digital micrometer (293-240, Mitutoyo, Japan). The diameter of the pin with a 30\u00b0 edge was 1004.5 um, and the diameter of the pin with the 45\u00b0 edge was 1005.0 urn. Holes with different diameters were fabricated in Al 6061 using a 700-f.!m drill and an 800-f.!m end mill. The hole depth was 2.3 mm. A smaller through-hole was fabricated inside each hole, as shown in Fig. 3, for air discharge. The hole diameters were measured with a digital microscope (SV32, Somethch, Korea) after they were filled with epoxy, mounted, and polished to a 200-f.!m depth from the plate surface. C. Insertion/Extraction Force Measurement The pins were inserted into a hole at 10 um/s to a 2-mm depth, and extracted at the same speed (see Table 2). Figure 6 shows the force data obtained during the insertion and extraction process. The force data were measured with the dynamometer at a frequency of 200 Hz. A low-pass filter of 50 Hz was used. The insertion and extraction took place over 200 s, and was performed three times per hole. The alignment between the pin and hole was guaranteed by swapping the end mill with the pin, thereby using the same coordinates for the hole fabrication and the pin-hole insertion. The pin was placed into the hole, and the insertion and extraction forces were measured." + ] + }, + { + "image_filename": "designv11_92_0000477_20080706-5-kr-1001.00289-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000477_20080706-5-kr-1001.00289-Figure2-1.png", + "caption": "Fig. 2. Range Sensor Image (a), Clearance Object (b), Environmental picture (c)", + "texts": [ + " Collision risk distance (dcollison) relies on the robot\u2019s and obstacle\u2019s speed. 2.3 Danger Index Danger index is defined to represent the collision risk, quantitatively. It is defined as the ratio of collision free speed area and collision speed area in a controllable speed area (Dynamic window) during the sampling frequency (Fox et al., 1997). Within the dynamic window speed range, the danger index becomes close to 1 when the collision area increases. On the other hand, the danger index is 0 when the robot is collision free. Fig. 2(a) shows the sensor data when a dynamic obstacle exists in front of a robot. The clearance object in the dynamic window is shown in Fig. 2(b). The collision time is calculated by considering the speed of the robot when obstacles are detected by the range sensor. The speed range of possible collision has 0 values. We will prove the motion safety from calculating the danger index. Fig. 3 shows the computed collision risk area without considering the robot\u2019s moving direction. However, it is not appropriate because we assume that collisions occur in front of the robot\u2019s heading. It does not consider collisions with obstacles behind the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000978_aim.2009.5229721-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000978_aim.2009.5229721-Figure3-1.png", + "caption": "Fig. 3. T2R1-type parallel manipulator with decoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b), branch with spatial motion of the moving platform (c), limb topology P R \u22a5 P \u22a5 R \u22a5 R-P \u22a5 P \u22a5\u22a5 R \u22a5 R -", + "texts": [ + " 1a, 2a and 3a the vector spaces iRGj have the following bases: (iRG1)=(v1, v2, v3, \u03c9\u03b1, \u03c9\u03b4), (iRG2)=(v1, v2, \u03c9\u03b1, \u03c9\u03b4), and (iRG3)=(v1, v2, v3, \u03c9\u03b1, \u03c9\u03b2, \u03c9\u03b4). In this configuration, Eq. (4) gives iSF=4 with (iRF)=(v1, v2, \u03c9\u03b1, \u03c9\u03b4). The moving platform has instantaneously four independent motions (two translations and two rotations) and the parallel mechanism has instantaneously the following structural parameters: iM=4, iN=1 and iT=0, given by (1)-(3). This constraint singularity occurs when q3=q2+ 2 2 2 1r q\u2212 for the solutions in Figs. 1 and 2, and q2=q3 for the solution in Fig. 3. In these configurations the rotation axes of the two last revolute joints of limbs G1 and G2 lay in the same plane. The axes of the last revolute joint of G1 limb and the before last revolute joint of G2 limb coincide. The parallel mechanism can get out from this constraint singularity by bifurcating in one of the two branches presented in Figs. 1-3b and c. The bifurcation in the constraint singularity can be used to change motion type of the moving platform 6. To achieve this change, one of the last revolute joints of G1 or G2-limb has to be instantaneously locked up when the moving platform passes through the constraint singularity", + " We note that the independent linear and angular velocities v1, v2, v3 and \u03c9\u03b1, \u03c9\u03b2, \u03c9\u03b4 of the characteristic point H have the directions parallel to the x0-, y0- and z0-axes of the reference frame. Equation (8) takes the following expressions: 1 1 1 2 22 2 2 1 3 1 2 2 1 11 2 1 1 0 0v q qv 1 0 q r q q q 1 1 r cos r cosr cos r q \u03d5\u03c9 \u03d5 \u03d5\u03d5 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5 \u23a1 \u23a4 \u23a1 \u23a4\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5= \u2212\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5\u2212\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5 \u23a3 \u23a6\u23a3 \u23a6 \u23a2 \u23a5 \u2212\u23a2 \u23a5 \u2212\u23a2 \u23a5\u23a3 \u23a6 (9) for the parallel manipulators in Figs. 1b,c and 2b,c, and 1 1 2 2 3 1 1 v 1 0 0 q v 0 1 0 q 1 1 q0 r cos r cos \u03d5\u03c9 \u03d5 \u03d5 \u23a1 \u23a4 \u23a2 \u23a5\u23a1 \u23a4 \u23a1 \u23a4\u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5= \u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5\u23a2 \u23a5\u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6\u23a3 \u23a6 \u23a2 \u23a5\u2212 \u23a2 \u23a5\u23a3 \u23a6 (10) for the solutions in Fig 3b,c. In Figs. 1-3, r1=HG=HK defines the dimensions of the moving platform. In (9) and (10), \u03d5 \u03b4= for the solutions with planar motion in Figs 1b, 2b and 3b, and \u03d5 \u03b1= for the solutions with spatial motion in Figs 1c, 2c and 3c. A new family of non overconstrained T2R1-type parallel manipulators with decoupled and bifurcated motion of the moving platform has been proposed. The approach has integrated the new formulae for mobility, connectivity, redundancy and overconstraint of parallel manipulators recently proposed by the author" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000087_12.819551-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000087_12.819551-Figure3-1.png", + "caption": "Fig 3 projection of tooth surface", + "texts": [ + " By the equation of meshing and the coordinate transformation, the pinion tooth surface is represented as follow: 1 1 p p( , )r r \u03b8 \u03c6= ur ur (4) Here: p\u03c6 is the rotation angle of cradle in the process of pinion generation. Proc. of SPIE Vol. 7130 71300F-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/26/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx By the criterion of Gleason, the length direction of tooth needs 9 measuring points and the height direction of tooth needs 5 measuring points. So 45 measuring points are measured by coordinate measuring machine and the corresponding coordinate can be gotten. These measuring points are showed by fig 3. The deviation of tooth surface is defined as the distance between actual tooth surface and theoretical tooth surface along normal of the measuring point on theoretical tooth surface. Data processing is based on the following procedure: Benchmark localization includes tilting localization and shaft localization[7]. The coordinate origin of the theoretical tooth surface is the vertex of pitch cones. In general, the shaft anchor of the actual tooth surface is located on the shaft collar. So the benchmarks of the theoretical tooth surface and the actual tooth surface should be coincident. The shaft localization is carried out by displacement along rotational axis of gear. The coordinate of measuring points of the actual tooth surface is expressed as r r r( , , )i i ix y z (i=1,\u2026,45). We define the third point along height direction and the fifth point along length direction as the datum mark, and its coordinate is expressed as rm rm rm( , , )x y z (Fig 3). The actual tooth surface is rotated by angle \u03c8 along its rotational axis, so that the datum mark on actual tooth surface is coincident with that on theoretical tooth surface. The rotational angle\u03c8 for tilting localization can be determined from the equation: rm rm rm rm rm ( , ) cos sin ( , ) sin cos ( , ) x x y y x y z z \u03b8 \u03c6 \u03c8 \u03c8 \u03b8 \u03c6 \u03c8 \u03c8 \u03b8 \u03c6 = +\u23a7 \u23aa = \u2212 +\u23a8 \u23aa =\u23a9 (5) Here: ),( \u03c6\u03b8x \u3001 ),( \u03c6\u03b8y \u3001 ),( \u03c6\u03b8z are the coordinate of datum mark on the theoretical tooth surface. All 45 measuring points are rotated by angle\u03c8 , and then tilting localization is completed. The coordinate of measuring points are expressed as: Proc. of SPIE Vol. 7130 71300F-3 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/26/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx ' r r ' r r ' r cos sin sin cos i i i i i i i i x x y y x y z z \u03c8 \u03c8 \u03c8 \u03c8 \u23a7 = + \u23aa = \u2212 +\u23a8 \u23aa =\u23a9 (i=1,\u2026,45) (6) Fig 3 shows the projection of tooth surface. All the measuring points are projected on the plane, which passes through the rotational axis of gear. The projection points of the actual tooth surface coincide with that of the theoretical tooth surface on this plane. The relationship between the measuring point of actual tooth surface and the corresponding theoretical point is represented by the following equations: 2 2 ' 2 ' 2 ' ( , ) ( , ) ( ) ( ) ( , ) i i i i i i i i i x y x y z z \u03b8 \u03c6 \u03b8 \u03c6 \u03b8 \u03c6 \u23a7 + = +\u23aa \u23a8 =\u23aa\u23a9 (i=1,\u2026,45) (7) By solving the equations, we can get the parameters of corresponding theoretical point ( , )i i\u03b8 \u03c6 , and then the corresponding theoretical point can be determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000640_icma.2009.5246170-FigureI-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000640_icma.2009.5246170-FigureI-1.png", + "caption": "Fig. I Model of walking machine .", + "texts": [ + " This method have realized at universal multibody simulation software FRUND. The essential of proposed approach is substitution redundant reactions forces to inertial forces. The main advantage of proposed method in comparison with existing methods is used the most common form of the equations of motion (1), that can be applied to system with arbitrary structure. Proposed theoretical method was incorporated at multibody dynamics simulation program FRUND. For numerical implementation was simulated four legs walking machine with two fingers on foots - fig. I. Model had following parameters: n =150, k =120, m =24, I =6 . Number of the rows of matrix Do are m + 6 x 4 = 48 and more of number degrees of freedom n - k = 30 . The cause of the static indefinite is three dimension contact force at every finger contact point. To solve one indefinite proposed include add big mass at contact point, linked with finger via ball joint. Redundant drivers may be transform similarly by include two or three parts kinematics chains. The first investigated regime of motion is the motion with a shift legs middle position to front as show in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003907_978-3-030-11434-3_26-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003907_978-3-030-11434-3_26-Figure3-1.png", + "caption": "Fig. 3. Absorber design for a tailpipe of a vehicle", + "texts": [ + " Figure 2b shows how the phase angle (between the input and output oscillation) versus r looks like when f = 0.1. Note regardless what f is, when r = 1, the phase angle is always = \u221290\u00b0. Typically, vibration absorbers can be designed to control system oscillations if the following criteria are met. (1) When the excitation frequency is constant. (2) When the excitation frequency is close to the system natural frequency. A common design of a vibration absorber for an exhaust pipe of a vehicle is shown below in Fig. 3. The absorber consists of a mass element, M2 mounted at the end of a stiffness element, K2, and a clamp to mount the absorber onto the system for vibration control. To understand why the excitation frequency needs to be close to the system natural frequency to work, it is useful to consider the phase angle information between the input and the absorber provided in Fig. 4. When the absorber mass is \u2212180\u00b0 out of phase with the input, cancelation will occur. Figure 5 shows the hand tremor displacement versus time graph for the patient model shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000364_jjap.47.1642-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000364_jjap.47.1642-Figure1-1.png", + "caption": "Fig. 1. Experimental setup for formation of birefringent films with twisted nematic structure in the PPLC film.", + "texts": [ + "38) The PPLC films were prepared by spin casting a solution of PPLC in methylene chloride on clear quartz slides without rubbing treatment, resulting in 1.6-mm-thick films. The PPLC films exhibited an amorphous structure upon deposition and the films showed good optical quality. Axis-selective photoreaction was performed using LP He-Cd laser beams, which emit a 325 nm light. From our previous observation, the absorption coefficient is about 4:3 104 cm 1 at a wavelength of 325 nm42) and this value is sufficiently high to absorb the incident light near the surface. Figure 1 shows the geometry of the exposure of writing beams adopted in this study. The LP He\u2013Cd laser beams were irradiated on both the film and substrate sides of the PPLC film independently in order to prevent light interference. The twisted angle ( ) between the polarization direction of the writing beam from the film side and that from the substrate side was varied between 0 and 90 . The induced twisted nematic structure was observed using a LP He\u2013Ne laser (633 nm) beam, which was incident normal to the film-side sample surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002894_ihmsc.2018.10168-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002894_ihmsc.2018.10168-Figure2-1.png", + "caption": "Figure 2. simplified model of height axis Because the experimental platform itself has a balance block to counteract the effective gravity component of the fuselage, the gravity component can be neglected and the differential equation of the height axis can be obtained", + "texts": [ + " Its three degrees of freedom are height axis, pitch axis and rotation axis. The kinetic model can be described by the equations of motion on these three axes. The power of the system is generated by the screw motor, and its thrust is approximately proportional to the voltage applied to it. Kf is the thrust constant of the combination of the motor helical propellers. V is the added voltage, and the thrust generated by the motor is (1) A. Height Axis Modeling According to the simplified model shown in Figure 2. the height axis is modeled in general, and the following dynamic differential equations are established: (2) Among them, J \u03b5 as the height of shaft inertia; \u03b5 high shaft deflection angle; K f motor constant; La body length to height of pivot points; p for the pitch axis deflection angle; V f andV b respectively before and after the voltage of the motor for the components of gravity. 270 978-1-5386-5836-9/18/$31.00 \u00a92018 IEEE DOI 10.1109/IHMSC.2018.10168 ( )bffa VVKpLJ += \u2022\u2022 cos\u03b5\u03b5 3 B. Pitching Axis Modeling According to the simplified model shown in Figure 3, the pitch axis movement by the difference between the two produced by the propeller thrust control, in the case that the current to the motor thrust is greater than the motor thrust, pitching axis Angle is positive, the helicopter fuselage would be positive pitch model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001780_978-3-658-21194-3_110-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001780_978-3-658-21194-3_110-Figure11-1.png", + "caption": "Figure 11: Principle sketch of the flow on a ground-effect car.", + "texts": [ + " Peter Wright and Tony Rudd at Lotus noted, in addition, something important during a wind tunnel campaign at Imperial Collage. A break down in the research equipment revealed that the curvature of the underbody was also significant for the downforce magnitude. Details of this experiment are vividly described in Ludvigsen (2012), and these findings resulted in the so called \u201cGround Effect\u201d which was used by most teams at the end of the seventies, and in the beginning of the eighties. One of the most beautiful \u201cground effect\u201d formula one cars was the Lotus 78 shown in figure 10. Figure 11 shows a typical underbody of a ground effect car. The curved side pods together with the blocking skirts are highlighted in the figure; the flow accelerated along the curved side pods and the low pressure zone created was enclosed by the side skirts to accomplish the downforce. Due to wearing of the skirts, against the road surface, a spring loaded movement acted on skirts to efficiently seal the gap throughout a full race, see Wright (2010). In connection to this, it could be remarked that it is highly questionable how these sliding skirts could be accepted in the formula one community since no movable devices were allowed at the time\u2026 The ground effects F1 cars were very efficient and launched many new thoughts, but among the drivers they were not popular, mostly because they were dangerous and the vehicles had extremely poor comfort" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure17-1.png", + "caption": "Fig. 17 Joint and link structure", + "texts": [ + " The cabin is also designed to be able to absorb the impact and protect a driver even if a large impact occurs by any chance, such as falling over or a collision with an obstacle. The robot has many electric components for driving including a battery, motor, and servo amplifier. For safety considerations, however, the high-voltage parts are completely isolated from the driver. The i-foot\u2019s legs are constructed so that each leg has 6 degrees of freedom, for a total of 12 degrees of freedom for the joints (Fig. 17). Each of the joint axes has a motor and harmonic drive reducer. The three hip joint axes and two ankle axes are orthogonal. The joint axis corresponding to the i-foot\u2019s knee bends in the opposite direction of human knees. This makes it possible to lower the cabin without interference with the legs, and the driver can mount and dismount without difficulty (Fig. 18). Figure 19 shows the i-foot interface for a driver. Figure 19a shows the left instrument panel and Fig. 19b shows the right joystick controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003764_sdpc.2018.8664965-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003764_sdpc.2018.8664965-Figure3-1.png", + "caption": "Fig. 3. Model sketch map of tapered roller bearing", + "texts": [ + " To facilitate calculation, the following assumptions are made without changing the bearing movement mechanism: \u25cf The bearing outer ring is fixed on the rigid element, with the speed 0, the lateral and vertical displacement; \u25cf The roller is arranged evenly on the raceway and is made of pure rolling; \u25cf contact stress is simplified and contact stiffness are the empirical values obtained experimentally. Lower part of bearing is the load area[23]. In the course of bearing movement, the total stiffness of bearings varies periodically, resulting in Varying Compliance(VC) vibration. Taking the right end bearing as an example, the schematic diagram of tapered roller bearing is shown in Fig. 3. The line velocity of the contact point of a roller core and inner ring is v1, the line velocity of contact point with outer ring is v2, the line speed of the rolling body is vc, the rotation angular velocity of the inner ring of bearing is w1, the rotation angle velocity of the bearing outer ring is w2, the inner ring is r, the inner diameter of the outer ring is R, the bearing clearance is C0, the bearing roller diameter is d, the bearing node diameter is D. Then v1, v1 are defined as follows: 1 1 2 2 v w r v w R (2) The cage line speed is given by 1 2 c 2 v v v (3) When v2=0, 1 1 c 2 2 v w r v (4) The angle speed of cage wc is defined as follows: 1 1 (1 cos ) 2 c d w w D (5) The rotation angle of the roller \u03b8i is defined as follows: c 0 0 2 ( 1) 1,2,3 ,i w t i i N N \uff0c (6) Where i=1,2,3,\u2026N0, N0 is the number of the balls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002082_1.4040813-Figure18-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002082_1.4040813-Figure18-1.png", + "caption": "Fig. 18 Pressure profiles versus oil supply flow rate, 10,000 rpm, 124 kPa specific load", + "texts": [ + " In both the flooded and starved models, decreasing oil supply flow rates result in decreasing bearing stiffness and damping. Increased predicted damping ratios with decreasing flow in both models suggest that the decreasing bearing stiffness is defining this trend. For all oil flow conditions, the starved model predicts lower stiffness and damping values than the flooded model across all speeds. Higher predicted damping ratios in the starved model suggest that the lower stiffness values are, once again, defining the trends seen in the resulting modal properties. Figure 18 illustrates the predicted pressure profiles of the starved bearing model for varying oil supply flow rates under the nominal load condition at 10,000 rpm. Similar to the speeddependent pressure profiles in Fig. 12, cavitation can be seen in the upper pads that increases in severity with decreasing oil supply flow rate, lessening the load on the lower pads. Again, a decrease in hydrodynamic film development in the upper pads and reduced loading of the lower pads result in the decreased bearing stiffness and damping seen in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000036_detc2009-87370-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000036_detc2009-87370-Figure4-1.png", + "caption": "Figure 4. Mathematical model for the nozzle geometry", + "texts": [ + " If a directed line, i pl , which is defined by (p, pn ), p i lateralS , intersects with any deposited layer Lj, ij , i.e., ijLl j i p (10) then collision between the nozzle and the deposited occurs. All the surface normals of the i lateralS need be tested. To be more accurate, the geometry of the nozzle needs be considered. The geometry and working distance of the nozzle are related to collision. Based on the physical process of deposition, a simplified geometry model of the nozzle is shown in Fig. 4. The geometry model of the nozzle consists of two parts: The laser beam Lbeam with limited beam length (equal to the working distance) beamH and the body of the nozzle, Lbody. Based on this geometric model, the collision between the nozzle and the deposited part can be detected using collision detection algorithms [13][14][15]. In nature, this method has the same constraints as the normal cutting method in multi-axis machining process. The difference between the deposition and the machining is that the deposition process can reduce or avoid the collision by optimizing the deposition order of layers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000952_etcs.2009.209-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000952_etcs.2009.209-Figure2-1.png", + "caption": "Figure 2 General 7R Series Robot", + "texts": [ + " The origin of world frame is located at the center of sphere (Fig.1). Z-axis is up in vertical direction through the origin and X-axis parallel to the axes of first joint in Robot 1.Thus Y-axis is obtained by right-hand screw rule. Longitudes and latitudes on the sphere are defined as geography. As shown in Fig.1, the equator is the intersection of the sphere and plane XOY in world frame, two Poles are the intersections of Z-axis and the sphere and Longitude 0o and 180o are the intersections of the sphere and plane XOZ in world frame. As shown in Fig.2, every local frame fixed with a link or a joint defined link frame or joint frame. The close-form equation is set up by D-H method. [ ] [ ] [ ] [ ] zjxwzrztcp 666000base RRTRT TTTI \u22c5\u22c5\u22c5\u22c5 \u22c5\u22c5\u22c5\u22c5\u22c5\u22c5\u22c5= \u03b3 \u03b1\u03b8\u03b1\u03b8 (1) Where [ ] \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 = 1000 100 0010 001 1000 0100 00 00 i i i ii ii i s a , cossin sincos T \u03b8\u03b8 \u03b8\u03b8 \u03b8 [ ] \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 = 1000 00 00 0001 ii ii i cossin sincos \u03b1\u03b1 \u03b1\u03b1 \u03b1 ),, 610( =i I \u2014 a 4\u00d74 identity matrix Tbase\u2014 a 4\u00d74 matrix of world frame to base frame [\u03b8i] \u2014 a 4\u00d74 rotation matrix of joint i around Zi-axis an angle \u03b8i (i=0,1,\u2026,6) Ti \u2014 a 4\u00d74 translation matrix of frame i to i+1 along Zi-axis a displacement si ,then along Xi-axis a displacement ai" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002890_j.matpr.2018.08.148-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002890_j.matpr.2018.08.148-Figure6-1.png", + "caption": "Fig. 6. Discrete model of the analysed spring.", + "texts": [ + " The spring in its middle part was reinforced with two stirrups connected with the yoke on the one side and with the rigid plate on the other. The stirrups were tightened with a torque of 500 Nm. Described mounting contributed to a 32% increase in stiffness of the spring. The experimental study [1-3] enable only spot measurement of the stresses and displacements. Complete overview of the results is possible through the numerical analysis. A computational model [4] of the examined spring was developed (fig. 6). It reflected all the necessary geometric features of the object. Due to the existing symmetry of the geometry and of the loading, the calculations were made only for a half of the model. Boundary conditions, enabling this type of the analysis of the structure were used. The loads acting between the leaves of the spring was considered by applying a proper contact connection. In order to verify a virtual model, comparison of the numerical and experimental tests was made. The authors compared obtained strains, displacements and the reaction forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000287_12.812031-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000287_12.812031-Figure1-1.png", + "caption": "Figure 1. The three standard planes of the heart, i.e. the long-axis, short-axis, and four chamber plane, are accessible through the parasternal, apical and subcostal scanning window in 2D echocardiography.", + "texts": [ + " In addition to providing road-map assistance, another advantage of this technology is that it may result in standardization while taking 2D views for measurement purposes. This is valuable in 2D US as the flexibility of the device causes high variability of measurement planes. Our approach is to use image registration to indicate, in a 3D model, how far/close the user is to a standard scan-plane. The feasibility of this technique is demonstrated in 2D echocardiography, where we apply 2D3D registration to guide the user to the standard planes (see Fig. 1). The method makes anatomy specific assumptions, but we perceive easy extension to other anatomy using an appropriate 3D model with extracted standard planes. In the past few years, there has been extensive work in the registration of US to other modalities (prominently CT and MR). Penney et al.1 and Blackall et al.2 have developed a method for registering 3D Freehand US to MR of the liver by using vascular landmarks. Recently, Wein et al.3\u20135 have proposed effective similarity measures for 3D Freehand US to CT registration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000741_apccas.2008.4746131-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000741_apccas.2008.4746131-Figure2-1.png", + "caption": "Fig. 2 shows the schematic diagram of the 3-RRRT parallel manipulator, where a11, a21, a31, in the moving platform, represent the centers of the T joints and a15, a25, a35 in the base platform, represent the centers of the R joints. \u25b3 a11a21a31 and \u25b3a15 a25 a35 are both regular triangles with circumcircle radius being r and R respectively. Three fixed coordinate system O-xiyizi i\u2208{1,2,3} has been established with the center of the base platform, with the oxi axis passing through the point ai1 respectively. and three moving coordinate system p-xiyizi has been attached to the moving platform in the center of the platform, with the pxi axis passing through the point ai5 respectively.", + "texts": [ + " In this paper, the inverse dynamics of 3-RRRT parallel manipulator is analyzed by a method based on the combination of theory of screws and the principle of virtual work, and numerical examples are given to show the characteristics of the methodology. II. DEFINITION OF MANIPULATOR STRUCTURE The 3-RRRT parallel manipulator shown in Fig. 1 consists of a moving platform, a fixed base, and three supporting limbs of identical kinematics structure. Each limb connects the moving platform to the fixed base by a tiger joint followed by three bars and three revolution joints. 745978-1-4244-2342-2/08/$25.00 \u00a92008 IEEE. Figure.2 Schematic diagram of 3-RRRT parallel manipulator For inverse position analysis, the position vector p=[px, py, pz]T of the reference point p in the moving platform is given and the matter is to find the limb angle of the input links. In the fixed coordinate system O-x1y1z1, according to the geometry of the manipulator, the following equations can be calculated for the first limb: 11 12 15e e e= = ; 13 14e e= ; 11 13 0e e =i (1) 1 11 11 12 11 12 13 13 11 12cos( ) cos( ) cos( )cos( )px R r L L L\u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8= \u2212 + + + + + (2) 1 13 13sin( )py L \u03b8= (3) 1 11 11 12 11 12 13 13 11 12sin( ) sin( ) cos( ) sin( )pz L L L\u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8= + + + + (4) Three bar angles of the first limb can be gotten from (2),(3),(4): ' 2 2 2 2 ' 1 1 11 12 13 1 11 ' 2 2 ' 2 2 11 1 1 1 1 ( ) sin( sin( ) 2 p p p p p p p x z L L L x a a L x z x z \u03c3 \u03b8 + + \u2212 + = \u2212 + + (5) ' 2 2 2 2 ' 1 1 12 13 11 1 12 11' 2 2 ' 2 2 12 13 1 1 1 1 ( ) sin( sin( ) 2( ) p p p p p p p x z L L L x a a L L x z x z \u03c3 \u03b8 \u03b8 \u03c3 + + + \u2212 = \u2212 \u2212 + + + (6) 1 13 13 sin( )py a L \u03b8 = (7) Where: 2 2 13 1 13 pL y L \u03c3 \u2212 = \u00b1 ; ' 1 1p px x R r= \u2212 + Similarly, angles of the bars of the second limb and the third limb can be calculated in the same way mentioned above" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000752_sice.2008.4655087-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000752_sice.2008.4655087-Figure4-1.png", + "caption": "Fig. 4 Inverted pendulum model as a simplified model of Tetsuro.", + "texts": [ + " In this study, \u03b8\u0307 ph j d was set to zero. The treadmill has two belts equipped with a DC motor for each, so that the speed of each belt (i.e., each leg) can be controlled independently. In different test stages, Tetsuro walks on the treadmill with the two belts either moving at the same speed (\u201ctied\u201d configuration) or different speed (\u201csplitbelt\u201d configuration). We can change the belt speed from 0 m/s to 0.6 m/s in every 0.005 m/s step manually. A 2-D inverted pendulum model such as that shown in Fig. 4 is constructed in order to plan the trajectory of a supporting leg in nonsteady state. In addition, we measured each joint angle in an absolute angle, not a relative angle. The state of efficient walking like human (normal subject) is motion in which the ankle torque output is approximately zero. The trajectory that satisfies these three equations on the nth step in a single step (0 \u2264 t \u2264 T0) can be expressed as follows. \u03c6(n)(t) = (1 + e\u2212aT0)eat \u2212 (1 + eaT0)e\u2212at (2) \ufffd a = \ufffd m1gl1 I1 \ufffd where I1 is the moment of inertia around the joint,m1 is the concentrated mass of link1, g is acceleration of gravity, l1 is the length of link1, and T0 is the supporting phase period" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002489_s13198-018-0747-4-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002489_s13198-018-0747-4-Figure3-1.png", + "caption": "Fig. 3 Structure sketch of equilibrium elbow", + "texts": [], + "surrounding_texts": [ + "According to the structural of the equilibrium elbow, we can get the normal stress r \u00bc M=Wx and shear stress s \u00bc T Wq. The M is the bending moment and T is the torque of the equilibrium elbow. Wx and Wq is the bending section coefficient and torque section coefficient. Based on the stress-strength interference theory, the state function of equilibrium elbow under different failure mode is: Static strength failure g1 \u00bc rs ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 \u00fe 3s2 p \u00f01\u00de rs is the static strength limit of the equilibrium elbow material Fatigue strength failure g2 \u00bc r 1K ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2 \u00fe 4 js\u00f0 \u00de2 q \u00f02\u00de r 1K is the fatigue strength limit, j is the shear stress equivalent coefficient Torsion stiffness failure g3 \u00bc h\u00bd h \u00bc h\u00bd Tl GWq \u00f03\u00de h is the steering angle, l is the length of the torsion shaft, G is the shear modulus The main failure modes of the equilibrium elbow are static strength failure, fatigue strength failure and torsion stiffness failure. Themain force that the component endured is the shock from the loadwheel, in this situation, the primary failuremodes is static strength failure. In the actual working situation of the tracked-vehicle, the equilibrium elbow has been endured the alternate loading as the complex road situation, the fatigue strength failure is also themain failuremode of the component. The equilibrium elbow is a combination component which the torsion force also can\u2019t be ignored, the torsion strength failure is the third failure modes of the component. The static strength failure usually occurs when an impact force beyond the strength limit loading on the component and made it cracked or fractured, the fatigue strength failure caused by the continuous impact loading on the component and the torsion strength failure caused by the twisting force loading on the component from the torsion shaft. These three failure modes are the main failure pattern of the equilibrium elbow by daily inspection. As anyone of the three failure modes occurred will make the equilibrium elbow failure, we consider the equilibrium as a series system, we can see the relationship of these three failure modes in Fig. 5." + ] + }, + { + "image_filename": "designv11_92_0000275_1.3273666-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000275_1.3273666-Figure11-1.png", + "caption": "FIGURE 11. Digitalized nozzle body.", + "texts": [ + " This attenuated energy would be the input parameter for the thermal model on the substrate in order to evaluate the melt pool geometry. Thus, a particle flow model has been developed to predict the powder concentration distribution in any point of the nozzle outlet. This model is based on the resolution of Navier-Stokes equations for turbulent regime using a k-s standard method. Once the mathematical problem is raised, it is necessary to define the geometry where this will be solved. Therefore, a full digitalization of the coaxial nozzle used in this work has been developed. Both real and digitalized nozzle body are shown in Fig. 11. Once the complete problem to be solved by FLUENT\u00ae is raised, only its simulation using the appropriate input parameters is necessary. The particle size is one of these parameters; generally given by its statistic distribution. The most used distribution for the laser cladding process simulation is the Rosin-Rammler distribution. This is unknown; so it is necessary to make a size study for each used material. Thus, the results obtained for W.Nr. 1.2379 powder are shown in Fig. 12, with these data it is possible the fitting to the Rosin-Rammler's distribution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.13-1.png", + "caption": "Fig. 11.13 Fiala model. Reproduced from Ref. [2] with the permission of Guranpuri-Shuppan", + "texts": [ + " The self-aligning torque due to the lateral displacement of the string v is given by M0 z \u00bc ks Za a vxdx\u00fe S a\u00fe r\u00f0 \u00de v1 v2\u00f0 \u00de=r: \u00f011:33\u00de The self-aligning torque due to the circumferential displacement u is given by M z \u00bc Cx Za a Zb 2 b 2 uydxdy; \u00f011:34\u00de where Cx is the shear spring rate of the tread per unit area in the circumferential direction. The total moment Mz is given by Mz \u00bc M0 z \u00feM z : \u00f011:35\u00de Suppose that the lateral displacement of the string in the contact patch v can be expressed by a straight line at a given slip angle a. Neglecting the circumferential displacement u, v is given by v \u00bc a a x\u00fe r\u00f0 \u00de: \u00f011:36\u00de The substitution of Eq. (11.36) into Eqs. (11.32) and (11.33) yields Fy \u00bc CFaa Mz \u00bc CMaa CFa \u00bc 2ks r\u00fe a\u00f0 \u00de2 CMa \u00bc 2ksa r r\u00fe a\u00f0 \u00de\u00fe a2=3 : \u00f011:37\u00de (1) Beam model and Fiala model Figure 11.13 shows the Fiala model [3], where D is the rigid wheel, B is the circular beam acting as the belt of a radial tire, C is the springs acting as the carcass, and T is the tread rubber. The circular beam has in-plane flexural rigidity EIz, where Iz is the moment of inertia of area, and E is Young\u2019s modulus of the belt in the circumferential direction. ks is the lateral fundamental spring rate per unit length in the circumferential direction discussed in Sects. 6.1.3 and 6.4.2. When the side force Fy is applied to the circular beam, the deformation of the circular beam can be solved as a bending problem of the beam on an elastic foundation as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002171_978-3-319-49574-3_12-Figure12.11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002171_978-3-319-49574-3_12-Figure12.11-1.png", + "caption": "Fig. 12.11 Polymer-housed arrester", + "texts": [ + " Different basic design principles are used for high-voltage arresters and mediumvoltage arresters. In the high-voltage field, mechanical requirements are much higher than in normal distribution applications. For this reason, porcelain housings are still used besides the growing number of composite hollow core insulators, so-called tube designs and directly molded designs. For distribution arresters in medium-voltage systems, porcelain housings have rapidly disappeared, and the directly molded design is used almost exclusively today (Fig. 12.11). The stack of MO resistor elements is mechanically supported by an internal cage structure, for example, made from FRP (fiberglass-reinforced plastic) rods. This insert is clamped between the end flanges with the help of compression springs. Additional supporting elements (not shown in the figure) may be necessary to fix the insert in the radial direction. What is important is the fact that this arrester, due to its enclosed gas volume, needs a sealing and pressure relief system. Depending on the voltage level, surge arresters can consist of either one single unit or alternatively several units" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.30-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.30-1.png", + "caption": "FIGURE 16.30 Illustration of the separation (divided) and integration (one-sheet) methods. (From Natsumi, F., Ikemoto, K., Sugiura, H., Yangfisawa, T., and Azuma, K., 1991, JSAE Review, Vol. 12, No. 3, pp. 58\u201363.)", + "texts": [ + " Offal is the scrap metal blank that falls off during sheet metal blanking from coil. Small pieces from the offal can be joined to form a functional single blank. 4. Reduced sealing needs. 5. Improved crash worthiness, since laser welds are stiffer than corresponding spot welds. The advantages of tailored blank welding are further illustrated in Table 16.5, which compares two traditional methods of producing the body side frame with tailored blank welding. In Table 16.5, the separation (divided) method uses individual components (Fig. 16.30), which are first formed and then joined, usually by spot welding. In the integration (one-sheet) method, cuts are made in a large sheet panel to leave one large panel of the desired shape, which is then formed. In a sense, tailored blank welding combines the two approaches by first welding appropriate individual components to form the large panel of the desired shape (Fig. 16.31). TABLE 16.5 Comparison of Two Traditional Methods of Producing Automotive Body Components with Components Produced by Tailored Blank Welding Divided Method One-Sheet Method Tailored Blanks Appearance Poor Good Good Accuracy Low High High Yield High (65%) Low (40%) High Material flexibility Selectable Fixed Selectable Number of dies required High Low Low Source: Table 1, Nakagawa et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003949_iscid.2018.10152-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003949_iscid.2018.10152-Figure4-1.png", + "caption": "Figure 4. Equivalent stress nephogram", + "texts": [ + " The pre tightening force of the expansion sleeve on the hub can be achieved through the contact of the rigid body to the drum. Therefore, the key point of the model is to realize load simulation, as shown in Figure 3. The circumference angle is 210 degrees. The load is uniform and the direction is perpendicular to the cylinder section. III. RESULT ANALYSIS The equivalent stress nephogram and the equivalent strain nephogram can be obtained by calculating the two steps with the finite element analysis software Abaqus. Figure 4 (a) is an equivalent stress nephogram of the drum under the load amplification of 500 times. For convenience of observation, the drum is cut symmetrically along the XOY plane to show the variation of stress inside the drum, as shown in Figure 4 (b). It can be clearly obtained from Figure 4 that the minimum stress of the roller is 4.251MPa and the maximum value is 112.8MPa. Higher stress is concentrated in the hub and the reinforcement ring, especially the maximum stress occurs just below the load of the inner ring of the reinforcement ring, which is also the monitoring point RF. Therefore, it is necessary to check the strength of the reinforcing ring. The reinforcing ring material is Q235, and its yield strength is 235 MPa. According to the safety requirements, the strength of the driving drum must be less than 70% of the material distinguishing strength, which is 164" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003825_1754337119831107-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003825_1754337119831107-Figure12-1.png", + "caption": "Figure 12. Predicted velocity fields (m/s) using computational fluid dynamics for two 4.5 mm pellets (2 and 3, Figure 2) at Re ~54,000 and Ma ~0.58.", + "texts": [ + " The free shear layer that left the pellet at the head rim formed an interface between the high-velocity freestream and a low-pressure recirculation region near the pellet surface between bands \u2018h\u2019 and \u2018n\u2019. Entrainment of fluid from this recirculation region into the layer maintained this region at low pressure. Close to the pellet tail at band \u2018n\u2019, the surface pressure begun to rise substantially. The pressure coefficient rose rapidly from approximately 20.6 at band \u2018n\u2019 to 20.1 at band \u2018r\u2019 on the pellet tail rim. This rise was caused by the shear layer impinging and reattaching onto the pellet surface just downstream of area band \u2018n\u2019. Figure 12 has been included in order to demonstrate that the rise in pressure observed experimentally in Figures 10 and 11 after band \u2018n\u2019 was due to flow reattachment. Figure 12 shows that the shear layer detaching at the head rim subsequently reattached onto the pellet surface near the tail, which agrees with the observations of Figures 10 and 11, showing that the pressure coefficient value rose after surface band \u2018n\u2019. Figure 12 was constructed for the dome and flat-head pellets using results obtained with computational fluid dynamics (CFD) simulation (OpenFOAM, free-stream velocity 200m/s, compressible rhoCentralFoam solver, RAS k-Omega-SST turbulence model, mesh 1.13 106 cells). Figure 12 also confirms the presence of a recirculation zone behind the pellet base. From Figure 10, it can be seen that the base of the pellet (bands \u2018s\u2019 to \u2018aa\u2019) lay within the rear recirculation region. As a consequence, the value of the pressure coefficient at the pellet base was approximately constant at 20.2. Figure 10 also shows the distribution of 27 local forces acting on the area bands \u2018a\u2019 to \u2018aa\u2019. All 27 forces shown in Figure 10 were pressure forces on the bands and these forces were calculated and resolved axially as described previously", + " Figure 15 suggests that the value of the pressure coefficient at the detachment point (between bands \u2018g\u2019 and \u2018h\u2019) affected the downstream location where the flow reattached onto the pellet tail surface. For example, Figure 15 shows that in the case of the dome-head pellet, reattachment occurred relatively early at area band \u2018n\u2019, while for the cavity-head pellet, reattachment occurred two area bands downstream at \u2018p\u2019. Therefore, Figure 15 suggests that reattachment and pressure recovery occurred earlier when the value of the coefficient at the upstream detachment point was low. A physical explanation for this observation can be proposed with the help of Figure 12. The free shear layer, detaching from the pellet head rim, is subjected to a radial pressure gradient because the free stream is at atmospheric pressure, while the forward recirculation zone is at lower, sub-atmospheric pressure. This causes the shear layer to curve towards the pellet axis and move closer to the pellet surface. As a result, the impingement point onto the rear pellet slope is also further upstream. This is what appears with the domehead pellet, where the lower pressure in the recirculation zone intensified the radial pressure gradient on the shear layer, increasing its curvature and causing it to impinge earlier onto the pellet sloped tail, around band \u2018n\u2019", + " The boundary layer detaching from the pellet tail (between bands \u2018q\u2019 and \u2018r\u2019) became a free shear layer, which engulfed a large, low-pressure, recirculation region behind the pellet base. As shown in Figure 15, the pressure within the recirculation region was sub-atmospheric, while that of the free stream was atmospheric. The free shear layer was therefore subjected to a radial pressure gradient, which caused it to curve towards the pellet axis and enclose the rear recirculation region, as shown in Figure 12. Figure 15 suggests that the pressure within this recirculation region had a positive relationship with the pressure at the tail detachment point. For example, the dome-head pellet had the highest coefficient value at area bands \u2018q\u2019 and \u2018r\u2019 and the highest average coefficient value for area bands \u2018s\u2019 to \u2018aa\u2019. Similarly, the cavity-head pellet had the lowest coefficient value at area bands \u2018q\u2019 and \u2018r\u2019 and lowest average value for area bands \u2018s\u2019 to \u2018aa\u2019. In fact, for the four pellets, a high positive correlation coefficient of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001649_012093-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001649_012093-Figure1-1.png", + "caption": "Figure 1. Start and end position of the manipulation object being mounted on the tool racks by android", + "texts": [ + " The motion synthesis along the 2 1234567890 \u2018\u2019\u201c\u201d specified paths of OL (of the arm) in the working envelop with consideration for object obstacles position leading to deadlocks can also be studied [10]. The most frequent movement task for an android is putting on a working object or taking it off the tool rack. Therefore, let us examine the structure of the knowledge base in virtual control of the android arm motion in this example. Suppose we need to model the motion of the arm mechanism to transfer the manipulation object from the point AS (the start point of the synthesized path) to the point AE (the end point) considering the location of tool racks P1 and P2 (figure 1). The position of the points \u0410S and \u0410E in figure 1 is assigned by their projections \u0410S 1, \u0410 S 2, \u0410 E 1 and \u0410E 2. In this figure subscript one or two respectively means that the points belong to the horizontal plane of projection (or the top view) or the frontal plane of projection (or the front view). The general view and the kinematic diagram of an android robot arm mechanism are represented in the paper [8]. Here, it is necessary to clarify whether the arm can reach the target point of the synthesized path using virtual modelling or not. In figure 1 the system of coordinates O1x1y1z1 is connected with the body of the android, the parameters xT, yT and zT defining the origin position of this system relative to the inertial coordinate system O0x0y0z0. The parameter xt determines the minimal assigned safe distance from the android body to the tool racks. The position of the reference point BC1 of the bottom tool rack is defined by the coordinates xB, yB and zB. The virtual modelling problem is solved using the knowledge base of the past experience", + " The parameters that define the working envelope projections with proper allowance made for the position of tool racks in this paper can be used by one of the component parts of the knowledge base about the past experience. Let us agree to use the first component of the knowledge base to calculate the optimal rest position (configuration) of the android arm relative to the manipulation object (with the account of the start point AH of the OL path). The optimal position in motion synthesis in the vector of velocities characterized by the parameters xs, ys (figure 1) and generalized coordinates qi specifies the position of the android arm and body relative to the point AS at which the solid angle Us obtained by the instantaneous states implementation takes the maximum value [8]. The specified position of the android robot is calculated for each certain plane of level \u0394i, assigned with a definite step of the coordinate z\u03941, z\u03942 and so on (figure 1), the curve of the function being plotted for this coordinate: Us = f(xs ,ys). (1) Let us consider the procedure to determine the graph of the function (1) for the plane \u03941. To determine the graph of the function (1) and array of configurations where the OL centre coincides with some points \u0410S of the plane \u03941, specified with a certain grid space, the starting point \u0410S \u03941 being prescribed. The parameters xS and yS determine the rest position of the start point \u0410S \u03941 \u03941 by the paths of the OL in the system O1x1y1z1, with the centre of the OL moving in the plane \u03941 (symbol defines that the point belongs to the geometric object \u03941). Figure 1 does not show the position of the point \u0410S \u03941. This position coincides with one of the rectangle vertices of the plane \u03941, its sides being determined by the parameters \u0394xS \u0438 \u0394yS . The set of point positions \u0410S \u03941, constructed within the specified interval \u0394xS and \u0394yS (figure 1) on the plane \u03941, is obtained by the arm motion synthesis with the range of motion minimization criterion [2, 10]. The specified grid step for definition of points \u0410S is determined by the modulus of the vector of linear velocity for the OL centre motion in the plane \u03941. The initial agreed value of the vector q\u0394i(q1, \u2026, q5) corresponding to the OL centre position at the point AS \u03941 provides the maximum angle Us value. This value is derived from the analysis of a number of arm configurations that provide the assigned values of parameters z1 and xt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003904_iiai-aai.2018.00132-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003904_iiai-aai.2018.00132-Figure1-1.png", + "caption": "Fig. 1. 2D inverted pendulum system", + "texts": [ + " The objective of this paper is to design and implement a 2D inverted pendulum control system using swarm based gain scheduling for PID and LQR controllers. The optimal gains of the controllers are searched by swarm algorithm using MATLAB with extracted parameters from the physical model. For each predefined angle within the range of \u00b15\u00b0 from the vertical equilibrium, a set of optimum gains is determined. Finally, we use PLC to match the real deflection angle with its corresponding set of controller gains. II. RELATED THEORIES Fig. 1(a) and Fig. 1(b) show the whole system of our 2D inverted pendulum prototype, and a two-link pendulum, respectively. The supporting cart of the 2D inverted pendulum is moveable along the horizontal axis. Besides, the pendulum, which is attached to a frictionless hinge, can move freely in This research was funded by Faculty of Technical Education, King Mongkut's University of Technology North Bangkok. Contract no. FTE-2561-01. 630 978-1-5386-7447-5/18/$31.00 \u00a92018 IEEE DOI 10.1109/IIAI-AAI.2018.00132 two dimension within the defined range" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-47-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-47-1.png", + "caption": "Fig. 19-47: M-n curves PM-motor", + "texts": [ + " With smaller power levels and especially servo applications closed loop chopper control can be encountered. The typical motors are then permanent magnet motors (PM-motor). Also independently excited machines may be used. This type of machine has just been dealt with so that we limit our study here to the PM-motor. With small DC-motors the magnetic field is usually produced by ceramic permanent magnets. It is clear that the characteristics of such motors are very similar to DC-motors with field windings. A PM-motor does have better properties: 1. practically linear M-n curves (fig. 19-47). The armature reaction has less influence on the flux than in the case of wound poles since: a) the permeability of ceramic material is extremely low (almost the same as air) b) the coercitive force of PM-material opposes change as a result of field and armature action. 2. small dimensions of the PM-motor. As a result of the high coercitive force of permanent magnets the radial dimensions of the poles are much smaller than for wound pole pieces which results in a PM-motor being smaller and lighter than a classic DC-motor of equivalent power 3. no power loss in the field winding 4. high starting torque. SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR 19.47 Remark: M-n curves (fig. 19-47) For a DC-motor we can write: E = k 1 \u0387 n \u0387 \u03a6 and M = k 2 \u0387 I a \u0387 \u03a6 ; I a = V a \u2212 E ______ R i = V a \u2212 k 1 \u0387 n \u0387 \u03a6 ____________ R i ; so that: M = k 2 \u0387 \u03a6 [ V a \u2212 k 1 \u0387 n \u0387 \u03a6 ____________ R i ] For a PM-motor \u03a6 is constant and therefore: M = k 3 \u0387 V a \u2212 k 4 \u0387 n 1. Torque in blocked state: With n = 0 is M = k 3 \u0387 V a = torque in blocked state (stalled torque) 2. No load speed: at M = 0 is n = k 3 ___ k 4 \u0387 V a = no load speed 3. The M-n curves: With a constant supply voltage V a and increasing speed n, torque M decreases linearly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001226_kikaic.74.2870-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001226_kikaic.74.2870-Figure2-1.png", + "caption": "Fig. 2 Schematic view of ball trajectory", + "texts": [], + "surrounding_texts": [ + "2870\n\u65e5\u672c\u6a5f\u68b0\u5b66 \u4f1a\u8ad6 \u6587\u96c6(C\u7de8)\n74\u5dfb748\u53f7(2008-12)\n\u8ad6 \u6587No.08-0501\n\u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b \u304a \u3051\u308b \u30ad \u30e5\u30fc\u306e\u885d \u7a81 \u7279\u6027 \u8a55 \u4fa1*\n(\u7403\u306e\u8ecc\u9053\u3092\u30ab\u30fc\u30d6\u3055\u305b\u308b\u5834\u5408)\n\u5cf6 \u6751 \u771f \u4ecb*1, \u9808 \u8cc0 \u4e00 \u535a*2 \u6c5f \u6fa4 \u826f \u5b5d*3, \u9752 \u6728 \u7e41*4\nNumerical Evaluation of Impact Characteristics of Cue in Billiards\n(Case Where Ball Trajectory is Curved)\nShinsuke SHIMAMURA*5, Kazuhiro SUGA, Yoshitaka EZAWA and Shigeru AOKI\n*5 Graduate School of Engineering, Toyo University, 2100 Kujirai, Kawagoe-shi, Saitama, 350-8585 Japan\nThe numerical evaluation of impact characteristics of a cue in billiards is performed for the case where the ball trajectory is curved by hitting a ball on its upper left or right part with an inclined cue. The effective numerical method developed in the previous papers, in which a cue and a ball are assumed to be isotropically elastic and rigid, respectively, is extended for this case. The extended method is verified with an experiment using a high speed camera. The result obtained with the extended method also shows that the curvature of the ball trajectory is large for the impact on its almost exactly left part with a slightly inclined cue, and this result agrees with an empirical one. It is found from a numerical simulation that if the Young's modulus or mass density of a shaft of a cue is large, the ball trajectory immediately after the impact is deviated largely and then is curved with a large curvature. Because the extended method can evaluate quantitatively this kind of effects of material properties, it is useful for a design of a cue.\nKey Words : Impact, Cue, Billiards, Curved Ball Trajectory, Efficient Numerical Method , Simulation\n1. \u7dd2 \u8a00\n\u30d3\u30ea\u30e4\u30fc \u30c9\u306f,\u30c6 \u30cb\u30b9\u3084\u30b4\u30eb \u30d5\u3068\u3068\u3082\u306b\u4eba\u6c17 \u304c\u9ad8 \u304f\n\u4e16 \u754c\u4e2d\u3067\u89aa \u3057\u307e\u308c\u3066 \u3044\u308b\u304c,\u30ad \u30e5\u30fc\u306e\u8a2d\u8a08\u306b\u95a2\u3059 \u308b\u7814\n\u7a76 \u306f\u5c11\u306a\u3044.\u30ad \u30e5\u30fc\u306f\u5148\u7aef\u304b \u3089\u30bf \u30c3\u30d7,\u5148 \u89d2(\u3055 \u304d\u3064 \u306e)\u304a \u3088\u3073\u30b7\u30e3\u30d5 \u30c8\u3067\u69cb\u6210\u3055\u308c,\u305d \u308c\u305e \u308c\u901a\u5e38,\u9769, \u30d7\u30e9\u30b9\u30c1\u30c3\u30af\u304a \u3088\u3073\u6728\u6750\u3067\u4f5c\u6210\u3055\u308c\u308b.\u30b7 \u30e3\u30d5 \u30c8\u306f\u524d\n\u534a\u5206 \u3068\u5f8c \u534a\u5206 \u306b\u5206\u304b\u308c,\u4e2d \u592e \u3092\u91d1\u5c5e\u88fd\u306e\u306d \u3058\u3067\u9023 \u7d50\u3055 \u308c\u308b \u3053\u3068\u304c\u591a \u3044.\n\u307e\u305f,\u30ad \u30e5\u30fc\u306f\u7af6\u6280\u306e\u5c40\u9762\u306b\u5fdc \u3058\u3066(1)\u7403 \u306e\u4e2d\u592e\u90e8 \u3092\u649e \u304f\u57fa \u672c\u7684\u306a\u649e \u304d\u65b9\u306e\u4ed6\u306b,(2)\u4e0a \u90e8\u307e\u305f\u306f\u4e0b\u90e8\u3092\n\u649e\u3044\u3066\u7403\u306b\u6b63\u56de\u8ee2 \u307e\u305f \u306f\u9006\u540c\u8ee2 \u3092\u4e0e\u3048\u308b\u649e\u304d\u65b9,(3)\n\u6a2a\u90e8\u3092\u649e\u3044\u3066\u7403\u306b\u30b9\u30d4\u30f3\u3092\u4e0e\u3048\u308b\u649e\u304d\u65b9,(4)\u659c \u3081\u4e0a \u90e8\u3092\u659c\u3081\u4e0b\u65b9\u306b\u649e\u3044\u3066\u7403\u306e\u8ecc\u9053\u3092\u30ab\u30fc\u30d6\u3055\u305b\u308b\u649e\u304d\u65b9 \u306a\u3069\u69d8\u3005\u306a\u649e\u304d\u65b9\u304c\u8981\u6c42\u3055\u308c\u308b.\n\u30ad\u30e5\u30fc\u3092\u5408\u7406\u7684\u306b\u8a2d\u8a08\u3059\u308b\u305f\u3081\u306b\u306f,\u30ad \u30e5\u30fc\u5404\u90e8\u306e \u5f62\u72b6,\u5bf8 \u6cd5\u304a\u3088\u3073\u6750\u8cea\u304c\u3053\u306e\u3088\u3046\u306a\u69d8\u3005\u306a\u649e\u304d\u65b9\u306b\u3088 \u308b\u885d\u6483\u529b,\u7403 \u306e\u904b\u52d5\u306a\u3069\u306e\u885d\u7a81\u7279\u6027\u306b\u4e0e\u3048\u308b\u5f71\u97ff\u3092\u5b9a \u91cf\u7684\u306b\u8a55\u4fa1\u3059\u308b\u5fc5\u8981\u304c\u3042\u308b.\u7b46 \u8005\u3089\u306f\u524d\u5831(1)\uff5e(3)\u3067, \u7403\u3092\u525b\u4f53,\u30ad \u30e5\u30fc\u3092\u5f3e\u6027\u4f53\u3068\u4eee\u5b9a\u3057,\u305d \u308c\u305e\u308c\u3092\u6709\u9650 \u5dee\u5206\u6cd5\u304a\u3088\u3073\u6709\u9650\u8981\u7d20\u6cd5\u3067\u89e3\u6790\u3057,\u4e21 \u8005\u3092\u9023\u6210\u3059\u308b\u3053 \u3068\u306b\u3088\u308a,\u4e0a\u8a18(1)\uff5e(3)\u306e \u649e\u304d\u65b9\u306b\u5bfe\u3059\u308b\u885d\u7a81\u7279 \u6027\u3092\u8a55\u4fa1\u3059\u308b\u7c21\u4fbf\u306a\u65b9\u6cd5(\u52d5 \u7684\u63a5\u89e6\u6709\u9650\u8981\u7d20\u6cd5\u306b\u6bd4\u3079 \u3066\u8a08\u7b97\u6642\u9593\u304c\u7d041/30)\u3092 \u958b\u767a\u3057\u305f.\n\u672c\u7814\u7a76\u3067\u306f\u524d\u5831\u3067\u958b\u767a\u3057\u305f\u65b9\u6cd5\u3092,\u4e0a\u8a18(4)\u306e \u649e\u304d\n\u65b9\u306b\u3082\u9069\u7528\u51fa\u6765\u308b\u3088\u3046\u306b\u62e1\u5f35\u3057,\u30b7 \u30e3\u30d5\u30c8\u306e\u6750\u8cea\u304c\u885d \u7a81\u7279\u6027\u306b\u4e0e\u3048\u308b\u5f71\u97ff\u3092\u660e\u3089\u304b\u306b\u3059\u308b.\u524d \u5831\u3068\u540c\u69d8\u306b, \u533a\u5225\u306e\u5fc5\u8981\u304c\u3042\u308b\u5834\u5408\u306b\u306f,\u53f0 \u306b\u5782\u76f4\u306a\u8ef8\u307e\u308f\u308a\u306e\u7403 \u306e\u56de\u8ee2\u3092\u30b9\u30d4\u30f3,\u305d \u306e\u4ed6\u306e\u8ef8\u307e\u308f\u308a\u306e\u56de\u8ee2\u3092\u8ee2\u304c\u308a\u3068\n\u547c\u3076\u3053\u3068\u306b\u3059\u308b.\n* \u539f\u7a3f\u53d7\u4ed82008\u5e746\u67082\u65e5. *1\n\u6771\u6d0b \u5927\u5b66 \u5927\u5b66 \u9662\u6a5f \u80fd \u30b7 \u30b9\u30c6 \u30e0\u5c02\u653b(\u3013350-8585\u5ddd \u8d8a \u5e02\u9be8 \u4e95\n2100). *2\n\u6771\u4eac\u7406 \u79d1\u5927 \u5b66\u7406\u5de5 \u5b66\u90e8(\u3013278-8510\u91ce \u7530\u5e02 \u5c71\u5d0e2641). *3 \u6b63 \u54e1,\u6771 \u6d0b\u5927 \u5b66 \u30b3 \u30f3 \u30d4\u30e5 \u30c6 \u30fc \u30b7 \u30e7\u30ca \u30eb \u5de5 \u5b66 \u79d1(\u3013350-8585\n\u5ddd\u8d8a \u5e02\u9be8\u4e952100). *4\n\u6771\u6d0b \u5927\u5b66 \u30b3 \u30f3\u30d4\u30e5\u30c6 \u30fc\u30b7 \u30e7\u30ca\u30eb\u5de5\u5b66 \u79d1.\nE-mail:ezawa@toyonet.toyo.ac.jp", + "\u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b\u304a \u3051 \u308b\u30ad\u30e5\u30fc \u306e\u885d\u7a81 \u7279\u6027 \u8a55 \u4fa1 2871\n2. \u89e3 \u6790\u65b9\u6cd5\u306e\u6982\u8981\n\u56f31\u306b \u793a\u3059\u3088 \u3046\u306b,\u7403 \u306e\u659c\u3081\u4e0a\u90e8 \u3092\u659c\u3081\u4e0b\u65b9\u306b\u649e \u3044\n\u3066,\u7403 \u306e\u8ecc\u9053\u3092\u30ab\u30fc\u30d6 \u3055\u305b\u308b\u5834\u5408\u3092\u8003\u3048\u308b.\u305f \u3060 \u3057,\n\u7403 \u3068\u5e8a\u306e\u63a5\u89e6\u529b\u306b\u3088\u308b\u4e21\u8005\u306e\u5909\u5f62\u306f\u7121\u8996\u3067\u304d\u308b\u307b \u3069\u5c0f \u3055\u3044\u3068\u4eee\u5b9a\u3059\u308b.\u524d \u5831(1)\uff5e(3)\u3068 \u540c\u69d8 \u306b,\u30ad \u30e5\u30fc\u5404\u90e8 \u3092\n\u7b49\u65b9\u5f3e\u6027\u4f53,\u7403 \u3068\u53f0\u3092\u525b\u4f53,\u30ad \u30e5\u30fc \u3068\u7403\u306e\u554f\u306e\u885d \u6483\u529b \u3092\u96c6\u4e2d\u8377\u91cd \u3068\u4eee\u5b9a\u3059\u308b.\n\u307e\u305a,\u885d \u6483\u529b \u3092\u65e2\u77e5 \u3068\u4eee\u5b9a \u3057\u3066,\u7403 \u306e\u904b\u52d5\u65b9\u7a0b \u5f0f\u3092\n\u6709\u9650\u5dee \u5206\u6cd5\u3067\u89e3\u304d,\u885d \u6483\u70b9\u306b\u304a\u3051\u308b\u7403\u306e\u5909\u4f4d \u3092\u8a08\u7b97\u3059 \u308b.\u307e \u305f,\u52d5 \u7684\u6709\u9650\u8981\u7d20\u6cd5\u306b\u3088 \u308a\u885d\u6483\u70b9\u306b\u304a\u3051\u308b\u30ad\u30e5 \u30fc\u306e\u5909\u4f4d \u3092\u6c42\u3081\u308b.\u6b21 \u306b,\u885d \u6483\u70b9\u306b\u304a\u3044\u3066\u3059\u3079 \u308a\u304c\u7121 \u3044 \u3068\u4eee\u5b9a \u3057\u3066,\u4e21 \u8005\u306e\u5909\u4f4d\u304c\u7b49 \u3057\u304f\u306a\u308b\u3088 \u3046\u306b\u9023\u6210\u3055\n\u305b\u3066,\u30b7 \u30f3\u30d7\u30ec\u30c3\u30af\u30b9\u6cd5(4)\u306b \u3088 \u308a\u885d\u6483 \u529b\u3092\u4fee\u6b63\u3059 \u308b. 3\u7ae0 \u3067\u7403\u306e\u904b\u52d5\u89e3\u6790 \u6cd5,4\u7ae0 \u3067\u30ad \u30e5\u30fc\u306e\u6709\u9650\u8981\u7d20\u89e3\u6790 \u6cd5, 5\u7ae0 \u3067\u4e21\u8005\u306e\u9023\u6210\u6cd5 \u3092\u8a73\u8ff0\u3059 \u308b.\n3. \u659c\u3081\u4e0a\u90e8\u3092\u649e\u3044\u305f\u5834\u5408 \u306e\u7403\u306e\u904b\u52d5\n3\u30fb1 \u904b\u52d5\u65b9\u7a0b \u5f0f \u56f31\u306b \u793a\u3059\u3088 \u3046\u306b,\u925b \u76f4\u4e0b\u65b9 \u3092y\u65b9 \u5411\u3068\u3057,\u30ad \u30e5\u30fc \u306b\u5782\u76f4\u306a\u65b9\u5411 \u3092z\u65b9 \u5411 \u3068\u3059 \u308b\u56fa \u5b9a\u76f4\u89d2\u5ea7\u6a19\u7cfbO-xy,z\u3092 \u63a1\u7528\u3059 \u308b.\u3053 \u306e\u5ea7\u6a19\u7cfb \u3092,\u539f \u70b9\u304c\u904b\u52d5\u3059 \u308b\u7403\u306e \u4e2d\u5fc3 \u306b\u306a \u308b\u3088\u3046\u306b\u5e73\u884c\u79fb\u52d5 \u3057\u305f\u5ea7\u6a19 \u7cfbO'-x'zy',z'\u3082\u7528\u3044\u308b.\u56f31\u306e \u89d2\u5ea6 \u03b1,\u03b2\u3067\u8868 \u3055\u308c \u308b\u7403 \u8868\u9762 \u306e\u70b9 \u3092\u649e \u304f\u5834\u5408 \u3092\u8003 \u3048\u308b.\u7403\u306e\u904b\u52d5\u65b9\u7a0b\u5f0f\u306f\u5f0f(1) \uff5e(6)\u3067 \u4e0e\u3048 \u3089\u308c\u308b.\n\u4e26\u9032\u904b\u52d5:\n(1)\n(2)\n(3)\n\u56de\u8ee2\u904b\u52d5:\n(4)\n(5)\n(6)\n\u305f\u3060 \u3057,\u03b1(t)\u304a\u3088\u3073 \u03b2(t)\u306f\u885d\u6483\u70b9 \u306e\u4f4d\u7f6e \u3092\u8868\u3059\u89d2\u5ea6\u3067\n\u3042 \u308a,\u56f31\u306e \u77e2\u5370\u306e\u65b9\u5411\u3092\u6b63\u3068\u3059 \u308b.b(t)\u306f \u7403 \u4e2d\u5fc3 \u306e \u5909\u4f4d\u30d9\u30af \u30c8\u30eb(\u7d30 \u5b57\u306f\u305d\u306e\u5927 \u304d\u3055,\u4ee5 \u4e0b \u540c\u69d8),t \u304a\u3088\u3073 \u30c9\u30c3 \u30c8(.)\u306f \u305d\u308c\u305e\u308c\u6642\u9593\u304a\u3088\u3073\u6642\u9593\u5fae\u5206, \u4e0b\u6dfb\u3048\u5b57x,y\u304a \u3088\u3073Z\u306f \u305d\u308c\u305e\u308cx,y\u304a \u3088\u3073Z\u65b9 \u5411 \u6210\u5206,f(t)\u306f \u885d\u6483 \u529b\u30d9 \u30af \u30c8\u30eb,Fr(t)\u306f \u6469\u64e6 \u529b\u30d9\u30af \u30c8\u30eb, Mr\u304a \u3088\u3073Ms\u306f \u305d\u308c\u305e\u308c\u8ee2\u304c \u308a\u304a\u3088\u3073\u30b9\u30d4\u30f3\u306b\u5bfe\u3059 \u308b \u6469\u64e6\u30e2\u30fc \u30e1\u30f3 \u30c8,\u03b8(t)\u306f\u7403\u306e\u56de\u8ee2\u89d2\u3092\u8868\u308f\u3059 \u30d9\u30af \u30c8\u30eb, M\u304a \u3088\u30730\u306f \u305d\u308c\u305e\u308c\u7403\u306e\u8cea\u91cf\u304a\u3088\u3073\u534a\u5f84\u3067\u3042\u308b.\u307e \u305f,\u56f32\u306b \u793a \u3059\u3088 \u3046\u306bS\u304a \u3088\u3073\u03c1\u306f\u305d\u308c\u305e\u308c\u7403\u3068\u53f0\u306e\n\u9593\u306e\u3059\u3079 \u308a\u901f\u5ea6\u30d9 \u30af \u30c8\u30eb\u304a\u3088\u3073\u56de\u8ee2\u306b\u3088\u308b\u7403 \u4e2d\u5fc3 \u306e\u901f \u5ea6 \u30d9\u30af \u30c8\u30eb\u3067 \u3042 \u308a,\u03b4(t)\u304a\u3088\u3073 \u03b3(t)\u306fS\u304a \u3088\u3073 \u03c1 \u3068x \u8ef8 \u306e\u306a\u3059 \u89d2\u5ea6 \u3067\u3042 \u308a,\u77e2 \u5370\u306e\u65b9\u5411 \u3092\u6b63 \u3068\u3059\u308b.\n\u7403\u4e2d\u5fc3\u306e\u901f\u5ea6\u30d9\u30af\u30c8\u30ebb,\u7403 \u306e\u89d2\u901f\u5ea6\u30d9\u30af\u30c8\u30eb\u03b8\u304a\n\u3088\u3073\u3059\u3079\u308a\u901f\u5ea6\u30d9\u30af\u30c8\u30ebS\u306e \u5404\u6210\u5206\u306e\u9593\u306b\u306f\u6b21\u5f0f\u304c \u6210\u308a\u7acb\u3063.\n(7)\n(8)", + "2872 \u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b\u304a \u3051\u308b\u30ad \u30e5\u30fc \u306e\u885d \u7a81 \u7279\u6027 \u8a55\u4fa1\n(9)\n\u56de\u8ee2\u306b\u3088\u308b\u7403\u4e2d\u5fc3\u306e\u901f\u5ea6\u30d9\u30af\u30c8\u30eb\u03c1 \u3068\u56de\u8ee2\u306e\u89d2\u901f\n\u5ea6\u30d9\u30af\u30c8\u30eb\u03b8\u306e\u5404\u6210\u5206\u306e\u9593\u306b\u306f\u6b21\u306e\u95a2\u4fc2\u304c\u3042\u308b.\n(10)\n(11)\n(12)\n(13)\n3\u30fb2 \u885d\u6483\u70b9\u3092\u8868\u308f\u3059\u89d2\u5ea6 \u885d\u6483\u70b9\u306e\u4f4d\u7f6e\u3092\u8868\u308f\n\u3059\u89d2\u5ea6 \u03b1(t),\u03b2(t)\u304a\u3088\u3073\u305d\u306e\u89d2\u901f\u5ea6a,\u03b2 \u3068\u56de\u8ee2\u306e\u89d2\u901f \u5ea6\u30d9\u30af\u30c8\u30eb\u03b8\u306e\u5404\u6210\u5206\u306b\u306f\u6b21\u306e\u95a2\u4fc2\u5f0f\u304c\u6210\u7acb\u3059\u308b.\n(14)\n(15)\n(16)\n\u5f0f(14)\uff5e(16)\u306e \u5de6 \u8fba \u306f \u56de\u8ee2 \u306b \u3088 \u308b,\u305d \u308c \u305e \u308cx,\ny'\u304a \u3088\u3073Z\u65b9 \u5411 \u306e\u885d \u6483 \u70b9 \u306e \u901f \u5ea6(\u6b63 \u78ba \u306b \u306f \u901f \u5ea6/R)\u3092\n\u03b1\u304a \u3088\u3073 \u03b2\u3067 \u8868 \u308f \u3057,\u53f3 \u8fba \u306f \u03b8x,\u03b8y\u304a \u3088\u3073 \u03b8z\u3067\u8868 \u308f \u3057\n\u3066\u3044 \u308b(\u4ed8 \u9332 \u53c2 \u7167),\u5f0f(15)\u306f,a\u3092 \u03b8x\u3068 \u03b8z\u3067\u8868 \u308f\n\u3057\u3066 \u3044 \u308b\u306e \u3067,\u3053 \u308c \u3092\u5f0f(14)\u307e \u305f \u306f(16)\u306b \u4ee3\u5165 \u3059\n\u308b \u3053 \u3068\u306b \u3088 \u308a,\u03b2 \u3092 \u03b8x,\u306e \u304a \u3088\u3073 \u03b8z\u3067\u8868 \u308f\u3059 \u3053 \u3068\u304c \u51fa\n\u6765 \u308b.\n3\u30fb3 \u89e3 \u6cd5 \u3042 \u308b\u6642 \u523bt\u306b \u304a \u3051 \u308bfx(t),fy(t)\u304a \u3088\u3073\n\u53cc')\u304c \u4e0e \u3048 \u3089\u308c \u305f\u5834 \u5408\u3092 \u8003 \u3048 \u308b.\u307e \u305a,S\u22600\u3068 \u4eee \u5b9a\u3059\n\u308b.Frx(t)\u3068Frz(t)\u306f \u5f0f(17)\u3068 (18)\u304b \u3089 \u03b4(t)\u306b\u3088 \u308a\u8868\n\u3055\u308c \u308b \u306e\u3067,\u6c42 \u3081 \u308b\u3079 \u304d\u5909\u6570 \u306fbx(t),by(t),bz(t),\u03b8x(t),\n\u03b8y(t),\u03b8z(t),s(t),sx(t),sz(t),\u03c1(t),\u03b4(t),\u03b3(t),\u03b1(t)\u304a \u3088\n\u3073 \u03b2(t)\u306e14\u500b \u3067 \u3042 \u308b.\nS\u22600\u306e \u5834 \u5408\n(17)\n(18)\n\u3053\u3053\u3067,\u03bc \u306f\u6469\u64e6 \u4fc2\u6570 \u3092\u8868 \u3057,\u7c21 \u5358 \u306e\u305f \u3081\u306b,\u52d5 \u6469\n\u64e6 \u4fc2\u6570 \u3068\u9759\u6469\u64e6 \u4fc2\u6570 \u3092\u533a\u5225\u305b\u305a\u306b0\u5b9a \u5024 \u3068\u3057\u3066\u3044 \u308b.g \u306f\u91cd\u529b\u52a0 \u901f\u5ea6 \u3067\u3042\u308b.\n14\u500b \u306e\u5909\u6570 \u306e\u5185,\u03b1(t)\u3068\u03b2(t)\u306f\u5f0f(15)\u3068(14)(\u307e\n\u305f\u306f(16))\u304b \u3089 \u03b8x(t),\u03b8y(t),\u304a\u3088\u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b. by(t)\u306f\u65e2\u77e5(by(t)=0)\u3067 \u3042\u308b.S(t),\u03c1(t),\u03b4(t)\u304a\u3088\u3073y(t)\n\u306f\u5f0f(10)\uff5e(13)\u304b \u3089sx(t),sz(t),\u03b8x(t),\u03b8y(t)\u304a\u3088 \u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b.sx(t)\u3068sz(t)\u306f \u5f0f(7)\u3068(9) \u304b \u3089bx(t),bz(t),\u03b8x(t)\u304a\u3088\u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b.\u3057 \u305f\u304c \u3063\u3066,5\u500b \u306e\u672a\u77e5\u6570bx(t),bz(t),\u03b8x(t),\u03b8y(t)\u304a\u3088\u3073 \u03b8z(t) \u306f\u5f0f(1),(3)\uff5e(6)\u306e5\u500b \u306e\u65b9\u7a0b\u5f0f \u304b \u3089\u89e3 \u304f\u3053\u3068 \u304c\u51fa\u6765 \u308b.\u521d \u671f\u6761\u4ef6 \u306f\u3053\u308c \u3089\u306e5\u500b \u306e\u5909\u6570\u304a \u3088\u3073\u305d\u306e\n\u6642\u9593\u5fae \u5206=0\u3067 \u3042\u308b.\u3053 \u308c \u3089\u306e\u5fae\u5206\u65b9\u7a0b\u5f0f \u306f\u6709\u9650\u5dee\u5206 \u6cd5\n(\u4f8b\u3048\u3070,Newmark\u306e \u03b2\u6cd5(\u03b2=1/6)(5))\u306b \u3088 \u308a\u89e3 \u304f\u3053\n\u3068\u304c\u51fa\u6765\u308b.\n\u3082\u3057,\u89e3 \u3044\u305f\u7d50\u679c\u6700\u521d\u306e\u4eee\u5b9a\u304c\u6210 \u7acb\u305b\u305a,S=0\u306b\n\u306a\u3063\u305f\u5834 \u5408\u306f,\u524d \u5831G)\u3068 \u540c\u69d8 \u306b\u5f0f(17)\u3068(18)\u306f \u6210 \u7acb \u3057\u306a\u3044\u305f\u3081,Frx(t)\u3068Frz(t)\u304c\u672a\u77e5\u6570 \u3068\u306a\u308b\u304c,\u7897)\u3068\n\u7897)\u304c0\u3067 \u3042\u308b\u304b \u3089,\u5f0f(7)\u3068(9)\u306b \u3088\u308abx(t)\u3068bz(t) \u306f \u03b8x(t)\u3068\u03b8z(t)\u3067\u8868 \u308f\u3055\u308c \u308b.\u3057 \u305f\u304c \u3063\u3066\u672a \u77e5\u6570 \u306e\u6570 \u3068 \u65b9\u7a0b\u5f0f \u306e\u6570\u304c0\u81f4 \u3057\u89e3 \u304f\u3053\u3068\u304c\u51fa\u6765 \u308b.\n\u306a\u304a,\u7d0410-3\u79d2 \u7a0b\u5ea6\u306e\u30ad\u30e5\u30fc \u3068\u7403 \u306e\u63a5\u89e6\u6642\u9593\u5185\u3067\na(t)=0\u3068 \u4eee\u5b9a(\u524d \u5831(1)\u306b \u3088 \u308a\u7403\u306e \u4e2d\u592e\u3092\u649e \u3044\u305f\u5834\u5408 \u306b \u306f\u63a5 \u89e6\u6642\u9593\u5185\u3067\u306f\u6b86 \u3069\u3059\u3079 \u308a\u306e\u307f\u3067\u56de\u8ee2\u306f\u7121\u8996\u3067 \u304d\n\u308b)\u3059 \u308b\u3068,\u4e0a \u8ff0\u306e\u65b9\u6cd5 \u306f\u524d\u5831(3)\u306e \u65b9\u6cd5 \u306b\u5e30\u7740\u3059 \u308b.\n4. \u4efb\u610f\u65b9 \u5411\u8377\u91cd \u3092\u53d7\u3051\u308b\u30ad\u30e5\u30fc\u306e\u5909\u4f4d\n\u56f31\u306b \u793a\u3059 \u3088\u3046\u306b,\u30ad \u30e5\u30fc \u3092\u659c \u3081\u4e0b\u65b9\u306b\u50be \u3051\u3066,(\u30ad\n\u30e5\u30fc\u306e\u8ef8\u65b9\u5411\u306b)\u7403 \u306e\u659c\u3081\u4e0a\u90e8\u3092\u649e \u304f\u5834\u5408 \u306b\u306f,\u30ad \u30e5 \u30fc\u306b\u306fx\n,y\u304a \u3088\u3073Z\u65b9 \u5411\u306e\u8377 \u91cdfx(t),fy(t)\u304a\u3088\u3073fz(t)\u304c\n\u4f5c\u7528\u3059\u308b.\u3053 \u306e\u5834\u5408\u306e\u885d\u6483\u70b9\u306b\u304a \u3051\u308b\u30ad\u30e5\u30fc\u306e\u5909\u4f4d \u3092 Cx(t),cy(t)\u304a\u3088\u3073cz(t)\u3068\u3057,\u3053 \u308c \u3089\u3092\u52d5\u7684\u5f3e\u6027 \u6709\u9650\u8981\u7d20 \u6cd5\u306b \u3088\u308a\u6c42 \u3081\u308b.\u89e3 \u6790\u306b\u306f,Nastran(6)\u3092 \u7528\u3044\u305f.\n\u88681\u306e \u89e3\u67901(Normal)\u6b04 \u306e\u6750\u6599 \u5b9a\u6570 \u3092\u6301\u3064,\u9577 \u3055 1.48m,\u5148 \u7aef\u90e8\u76f4\u5f840.013m,\u5f8c \u7aef\u90e8\u76f4\u5f840.032m\u306e \u30ad\u30e5 \u30fc\u3092\u8003 \u3048\u308b.\u56f33(a)\u306b \u30ad\u30e5\u30fc\u306e\u30e1 \u30c3\u30b7\u30e5\u5206\u5272\u56f3 \u3092\u793a\n\u3059.\u56f33(b)\u306b \u5185\u90e8\u306e\u69cb\u9020 \u3092\u793a\u3059\u305f\u3081\u306b\u65ad\u9762\u56f3\u3092\u793a\u3059. 8\u7bc0 \u70b9\u306e6\u9762 \u4f53\u8981\u7d20\u307e\u305f\u306f6\u7bc0 \u70b9\u306e5\u9762 \u4f53\u8981\u7d20(\u5168 \u8981 \u7d20\u65706363\u500b,\u5168 \u7bc0\u70b9\u65706184\u500b)\u3092 \u7528 \u3044\u3066\u5206\u5272 \u3057\u3066 \u3044\u308b.\u306a \u304a,\u306f \u308a\u67f1\u8981\u7d20 \u3092\u7528\u3044 \u308b\u3068\u3088 \u308a\u7c21\u4fbf \u306a\u65b9\u6cd5\u304c\n\u5f97 \u3089\u308c \u308b\u3068\u601d\u308f\u308c \u308b\u304c,\u4eca \u5f8c\u306e\u8ab2\u984c \u3068\u3057\u305f\u3044.\n\u885d\u6483\u529b \u3092\u96c6 \u4e2d\u8377\u91cd \u3068\u3057\u305f\u305f\u3081\u306b\u885d \u6483\u70b9 \u306e\u5909\u4f4d\u306b\u8aa4\n\u5dee\u304c\u751f \u3058\u308b\u304a\u305d\u308c \u304c\u3042\u308b\u306e\u3067,\u3053 \u308c \u3092\u5c0f \u3055\u304f\u3059 \u308b\u305f\u3081 \u306b,\u30d8 \u30eb \u30c4\u306e\u63a5\u89e6\u7406\u8ad6(7)\u3067\u4e88\u60f3 \u3055\u308c \u308b\u63a5\u89e6\u534a\u5f84\u306b\u5408 \u308f \u305b\u3066\u885d\u6483\u70b9\u4ed8\u8fd1 \u306e\u8981\u7d20\u306e\u5927\u304d \u3055\u3092\u6c7a\u5b9a \u3057\u305f.\u3059 \u306a\u308f\u3061 \u4e88\u60f3\u6700\u5927\u63a5\u89e6 \u534a\u5f84(\u304a \u3088\u305d2mm)\u306e1/2\u500d \u306e1mm\u3068" + ] + }, + { + "image_filename": "designv11_92_0002206_s1068798x18070043-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002206_s1068798x18070043-Figure10-1.png", + "caption": "Fig. 10. Moments in the cutting process: (a) steady cutting; (b) beginning of transition from steady cutting to tool exit; (c) end of transition; (d) tool exit from the blank; (e) end of cutter\u2013chip contact; (f) complete cutter exit from blank.", + "texts": [ + " For reasons of space, we present only the formulas for cutting in the exit zone that are required for the calculations. For more complete theoretical understanding, it is expedient to review the theory for the steady cutting zone [1\u201314]. In steady cutting, the main characteristics (the chip shrinkage, inclination of the shear plane, cutting force, and temperature) are constant or vary very slowly [15]. An important difference between the steady cutting zone and the tool\u2019s exit zone is that undeformed metal is constantly supplied to the zone of intense plastic deformation in the former case (Fig. 10a, zone OABC). In tool motion to the free end of the blank, the zone of intense plastic deformation (OABC) reaches the free surface of the blank at some instant (Fig. 10b). This cutter position is regarded as the onset of the transition from the steady cutting zone to the tool\u2019s exit zone. Thereafter, the supply of undeformed metal to zone OABC gradually decreases to zero. At the same time, this zone shrinks. When the supply of undeformed metal is over, the transition zone ends (Fig. 10c), and the exit zone begins (Fig. 10d). When the supply of undeformed metal to zone OABC is discontinued, it shrinks, while the cutter continues to move toward its exit from the blank. Since crack formation along the conditional shear plane is typical of cutting, the chip will separate from the front surface when the cutter is at some distance from the end of the black, on account of the crack that forms (Fig. 10e). As a result, the forces between the cutter and the blank will be due to contact of the machined surface with the tool\u2019s rear surface, and cutting will stop (Fig. 10f). We now consider the cutting process at the edge in the moment of chip breakaway from the blank and the cutter\u2019s front surface. The cutting scheme for this moment is shown in detail in Fig. 11, which is valid for both positive and negative values of the rake angle \u03b3. To simplify the analysis, we assume a plane strain state. In other words, we assume zero strain perpendicular to the plane of the drawing. In addition, we make the following assumptions. The cutter is introduced in the blank at speed v0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003990_j.enganabound.2019.04.008-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003990_j.enganabound.2019.04.008-Figure2-1.png", + "caption": "Fig. 2. The sketch map of discrete inner race.", + "texts": [ + " For the contact problems with friction, the incremental loading ethod is adopted in this paper and the boundary integral equation s expressed as follows [18] : \ud835\udc36 \ud835\udc58 \ud835\udc56\ud835\udc57 ( \ud835\udc4b )\u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a \ud835\udc57 ( \ud835\udc4b ) + \u222b\u0393\ud835\udc58 \u2212\u0393\ud835\udc58 \ud835\udc36 \ud835\udc47 \u2217 \ud835\udc56\ud835\udc57 ( \ud835\udc4b, \ud835\udc4c ) \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a \ud835\udc57 ( \ud835\udc4c ) \ud835\udc51\u0393 + \u222b\u0393\ud835\udc58 \ud835\udc36 \ud835\udc47 \u2217 \ud835\udc56\ud835\udc57 ( \ud835\udc4b, \ud835\udc4c ) \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a \ud835\udc57 ( \ud835\udc4c ) \ud835\udc51\u0393 = \u222b\u0393\ud835\udc58 \u2212\u0393\ud835\udc58 \ud835\udc36 \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 ( \ud835\udc4b, \ud835\udc4c ) \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a \ud835\udc57 ( \ud835\udc4c ) \ud835\udc51\u0393 + \u222b\u0393\ud835\udc58 \ud835\udc36 \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 ( \ud835\udc4b, \ud835\udc4c ) \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a \ud835\udc57 ( \ud835\udc4c ) \ud835\udc51\u0393 (1) here k represents object A or B, X represents source point, Y represents eld point, C ij ( X ) represents the coefficient of geometrical variation at he source point X , \u0394\ud835\udc62 \ud835\udc58,\ud835\udc5a \ud835\udc57 and \u0394\ud835\udc61 \ud835\udc58,\ud835\udc5a \ud835\udc57 represent the displacement and tracion of k -object in the j direction under the global coordinate system espectively when the m th step increment is loaded, and \ud835\udc48 \u2217 \ud835\udc56\ud835\udc57 ( \ud835\udc4b, \ud835\udc4c ) and \u2217 \ud835\udc56\ud835\udc57 ( \ud835\udc4b, \ud835\udc4c ) represent the displacement and traction fundamental solutions f elastic problem respectively. .2. Bearing boundary element Considering the traction characteristics of roller bearing, bearing oundary elements are used to simulate the bearing contact elements. ig. 2 shows the discrete model of inner race and the position relationhip with rollers. From Fig. 2 we know, one bearing boundary element is divided into wo sub-elements, and there is continuous traction on sub-element \u03931 \ud835\udc56 nd no traction on sub-element \u03932 \ud835\udc56 . The normal traction on the sublement \u03931 \ud835\udc56 is presented as linear distribution along the width direction nd the length direction [12] . As shown in Fig. 3 , there are two kinds f bearing boundary elements, BBE1 and BBE2. Considering the effect of friction and the incremental load, the in- remental traction on sub-element \u03931 \ud835\udc56 is expressed as follows: \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56, I \ud835\udf093 = 4 \u2211 \ud835\udefd=3 \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56,\ud835\udefd \ud835\udf093 (1 + (\u22121) \ud835\udefd\ud835\udf091 )( \ud835\udf092 \u2212 \ud835\udf090 2 )\u2215(1 \u2212 \ud835\udf090 2 ) \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56, II \ud835\udf093 = 2 \u2211 \ud835\udefd=1 \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56,\ud835\udefd \ud835\udf093 (1 + (\u22121) \ud835\udefd+1 \ud835\udf091 )( \ud835\udf092 \u2212 \ud835\udf090 2 )\u2215(1 + \ud835\udf090 2 ) (2) here, I and II represent the type of bearing boundary elements BBE1 nd BBE2, respectively, \ud835\udf091 , \ud835\udf092 and \ud835\udf093 represent the directions under local oordinate system, \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56, I \ud835\udf093 and \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56, II \ud835\udf093 represent the increment normal ractions of sub-element \u03931 \ud835\udc56 in BBE1 and BBE2, respectively, when the th step increment is loaded, \u0394\ud835\udc61 \ud835\udc5a,\ud835\udc56\ud835\udefd \ud835\udf093 represents the increment normal raction of the \ud835\udefdth node in element I , when the m th step increment is oaded, and \ud835\udf090 2 represents the coordinate values of inside edge of sub- lement \u03931 \ud835\udc56 , in the direction of \ud835\udf092 ( 1 \u2212 |\ud835\udf090 2 | is half-width of contact area)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002198_978-3-319-99270-9_25-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002198_978-3-319-99270-9_25-Figure1-1.png", + "caption": "Fig. 1. Rotor-bearing-seal system of a gas turbine.", + "texts": [ + " In this paper, the authors adopted the seal force model of brush seal and the nonlinear oil-film force model based on short bearing theory considering the lateral deflection of the disks, in order to build the nonlinear dynamic model of a multi-disk rotor-bearing-brush seal system. The effects of the rotor speed and eccentricity phase difference on the vibration response and dynamic behavior of a multi-disk rotorbearing-seal system were discussed under different operating conditions by axis orbit, Poincar\u00e9 map, and spectrum cascade. 2 Nonlinear Dynamic Model of a Multi-disk Rotor-BearingBrush Seal System 2.1 Nonlinear Dynamic Model of the Multi-disk Rotor-Bearing-Seal System Figure 1 shows the rotor-bearing-seal system of a gas turbine. The finite element model is obtained by discretization based on the structural features of the system, as shown in Fig. 1a. In this paper, the compressor and turbine are simplified as disk m8 and disk m9, which located at joint 8 and joint 9, respectively. Similarly, the supporting bearings are simplified as disk m4 and disk m12, which located at joint 4 and joint 12, respectively. Considering the lateral deflection of the disks, the nonlinear dynamic equation of the system can be obtained by the simplified model as below: M\u20acq\u00feC _q\u00feKq \u00bc Fg \u00feFb \u00feFs \u00feFe \u00f01\u00de With C \u00bc Kq \u00bc 6EI 1 l348 2x8 \u00fe 2x4 \u00fe hy8l48 \u00fe hy4l48 h i 6EI 1 l348 2y8 \u00fe 2y4 \u00fe hx8l48 \u00fe hx4l48\u00f0 \u00de h i 6EI 1 l348 2x8 \u00fe 2x4 \u00fe hy8l48 \u00fe hy4l48 \u00fe 1 l389 2x9 \u00fe 2x8 \u00fe hy9l89 \u00fe hy8l89 h i 6EI 1 l348 2y8 \u00fe 2y4 \u00fe hx8l48 \u00fe hx4l48\u00f0 \u00de\u00fe 1 l389 2y9 \u00fe 2y8 \u00fe hx9l89 \u00fe hx8l89\u00f0 \u00de h i 2EI 1 l289 3y9 3y8 hx9l89 2hx8l89\u00f0 \u00de 1 l248 3y8 3y4 hx8l48 2hx4l48\u00f0 \u00de h i 2EI 1 l289 3x9 3x8 hy9l89 2hy8l89 1 l248 3x8 3x4 hy8l48 2hy4l48 h i 6EI 1 l389 2x9 \u00fe 2x8 \u00fe hy9l89 \u00fe hy8l89 \u00fe 1 l3912 2x12 \u00fe 2x9 \u00fe hy12l912 \u00fe hy9l912 h i 6EI 1 l389 2y9 \u00fe 2y8 \u00fe hx9l89 \u00fe hx8l89\u00f0 \u00de\u00fe 1 l3912 2y12 \u00fe 2y9 \u00fe hx12l912 \u00fe hx9l912\u00f0 \u00de h i 2EI 1 l2912 3y12 3y9 hx12l912 2hx9l912\u00f0 \u00de 1 l289 3y9 3y8 hx9l89 2hx8l89\u00f0 \u00de h i 2EI 1 l2912 3x12 3x9 hy12l912 2hy9l912 1 l289 3x9 3x8 hy9l89 2hy8l89 h i 6EI 1 l3912 2x12 \u00fe 2x9 \u00fe hy12l912 \u00fe hy9l912 h i 6EI 1 l3912 2y12 \u00fe 2y9 \u00fe hx12l912 \u00fe hx9l912\u00f0 \u00de h i 2 66666666666666666666666666666666666666666664 3 77777777777777777777777777777777777777777775 where M is the mass matrix of the system, M \u00bc Mx 0 0 My , Mx \u00bc My \u00bc diag \u00bdm4;m8; Jd8;m9; Jd9;m12 , C is the damping matrix of the system, C \u00bc Cx 0 0 Cy , Cx \u00bc Cy \u00bc diag\u00bdc4; c8; ch8; c9; ch9; c12 , K is the stiffness matrix of the system, q is the displacement of geometry center Oi in the X and Y direction, respectively, q \u00bc \u00bdx4; y4; x8; y8; hx8; hy8; x9; y9; hx9; hy9; x12; y12 T , Fg is the gravity vector of the system, Fg \u00bc \u00bd0;m4g; 0;m8g; 0; 0; 0;m9g; 0; 0; 0;m12g T , Fb is the nonlinear oil-film force vector [8, 9], Fb \u00bc \u00bdFbx4;Fby4; 0; 0; 0; 0; 0; 0; 0; 0;Fbx12;Fby12 T , Fbxi Fbyi \" # \u00bc S0 fbxi fbyi \" # , i \u00bc 4; 12, Fs is the seal force vector [10], Fs \u00bc \u00bd0; 0;Fsx8;Fsy8; 0; 0; Fsx9;Fsy9; 0; 0; 0; 0 T , Fsxi Fsyi \" # \u00bc Fbi cos\u00f0a\u00fe h l /\u00de Fbi sin\u00f0a\u00fe h l /\u00de \" # , i \u00bc 8; 9, Fe is the unbal- anced force vector, Fe \u00bc \u00bd0; 0;m8eu8x 2 cos\u00f0xt\u00de;m8eu8x 2 sin\u00f0xt\u00de; 0; 0;m9eu9x 2 cos\u00f0xt\u00fe r\u00de; m8eu8x 2 sin\u00f0xt\u00fe r\u00de; 0; 0; 0; 0 T : For the facility of calculation, dimensionless transformations are introduced into the Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003016_j.compstruct.2018.11.095-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003016_j.compstruct.2018.11.095-Figure3-1.png", + "caption": "Fig. 3. Surface curvatures: (a) CAD of the fuselage section with k min1, and k max2, at a surface point (b) Drawings of the tape trajectory generated with the geometryinduced mechanism on a double-curved surface.", + "texts": [ + " However, the available commercial software solutions are not sufficient, due to our specific set of constraints. In this section we introduce the calculation of our tape laying process. Two factors must be considered for a wide tape application on doubly-curved surfaces: firstly, the curvature of the surface geometry and secondly the material draping properties. To characterise the curvature of the free-form surface we use the criteria of the Gaussian curvature G [20]. For the fuselage section we focus on a positive curved surface (see Fig. 3(a)). The Gaussian curvature can be calculated for each surface point, as the product of the minimum k min1, and maximum k max2, geodetic curvature of the surface. Application of material on doubled-curved surfaces can lead to deformations of the material (e.g. wrinkles, buckling). Therefore, the second factor affecting the appearance of wrinkles are the material draping properties, especially the bending-stiffness and shear-stiffne ss [21]. With our application mechanism we aim a wrinkle-free tape application", + " In contrast to the steering mechanism, tape edge lengths ( =L Ll r) are equal. In contrast to the shearing mechanism, the direction in which it is applied is the same as that of the tape\u2019s orientation. Less variations in material width W occur here. These conditions basically enable to apply wider tapes using an ATL process. However, the tape trajectory (Ll = Lr) on a double-curved surface must be generated. The tape trajectory for equal edge lengths cannot be explicitly calculated on a free-form surface. Fig. 3(b) shows a sketch of the generation elements for the trajectory on a doubly curved surface (1). The algorithm\u2019s basic procedure for generating a tape trajectory with a given starting point (2) and direction (3) on a surface is as follows: we first choose the starting point (2) as the centre of the lower edge of the tape. We then compute the remaining lower edge by constructing two geodesic curves (4,5) on the surface, in the two directions orthogonal to the given starting direction. The first point of the left tape edge (L0) and the first point of the right tape edge (R0) are then obtained for half of the given tape width WT ", + " Subsequently, an application-driven fitness criterion is defined and implemented as a two-phase iterated local search. Fig. 5 displays two exemplary multiple tape designs to describe the constraints and parameters of the optimisation. A multiple tape design for a surface (1) has to fulfill three constraints: \u2022 The surface (1) must be completely covered (However, tapes can start and end on an enlarged surface (7) for calculation reasons) \u2022 The tape trajectories must be generated with the geometry induced mechanism (equal edge lengths, see Fig. 3(b)) \u2022 A continuous overlap distance orthogonal to the application direction has to be ensured between aligned tapes (2). The overlap must be between a minimum distance lapmin and maximum distance lapmax Additionally, the following parameters can be varied and thus describe the room for optimisation: The start point location of the tapes (3), the starting direction of the tapes start (4), the tape width WT (5), the tape length LT (6) and the number of tapes NT . A reasonable limitation of the parameters and the focus on one fitness criterion is beneficial. For our use case, the limitations have mainly been driven by considering both geometric and automation issues. The maximum kmax and the minimum kmin curvature of typical double-curved fuselage section differ greatly. The maximum curvature kmax is orientated in an almost circumferential direction, thus orientated almost in the yz-plane (see Fig. 3(a)). As a low deformation of the compaction roller is advantageous for the tape laying process, we chose to use the minimum deviation of the tape orientation dev from the yzplane as a reasonable fitness criterion. Since the fuselage section is an aerodynamic surface, curvature variations are low and smooth. As such, we assume that aligned tape orientations for reasonable multiple tape designs are approximately parallel. We implement that limitation in our optimisation as we initially suggest the same orientation for aligned tapes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002523_ilt-11-2017-0351-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002523_ilt-11-2017-0351-Figure6-1.png", + "caption": "Figure 6 Internal schematic diagram of the body of the test rig", + "texts": [ + " However, an increase in the velocity gradient can result in the angle of the pad saltate increasing which can damage the pad. Therefore, in order to improve the life of the bearing as much as possible, the velocity gradient of the speed curve should choose a balanced value. 4. Verification experiment To verify the numerical model, the experiment was executed with a test rig for the water lubricated tilting-pad thrust bearing (Zhang et al., 2017), as shown in Figure 5. Test cavity include torque sensor, loading device, thrust plate and tilting-pad thrust bearing, as shown in Figure 6. The torque sensor was used to measure the friction torque which would affect the performance and life of the tilting-pad thrust bearing. The error of the system friction torque caused by the friction torque of the ball bearing, which should be reduced in the measurement results, can be evaluated before the test. The loading device was a hydraulic cylinder connected to a hydraulic station to provide a constant load to the thrust pad. A pair of symmetric tilting-pad thrust bearings was arranged face to face to balance the axial force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003583_scems.2018.8624797-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003583_scems.2018.8624797-Figure2-1.png", + "caption": "Fig. 2. Relative permeability distribution", + "texts": [ + " More importantly, if the saturation of magnetic path is taken into consideration, the distribution of solid rotor\u2019s relative permeability will change in a dynamic process. For example, at the beginning of a dynamic process, the solid rotor will generate a large eddy current to maintain the initial flux linkage, at this time, the eddy current will concentrate on the surface of the solid rotor, hindering flux linkages that pass through air-gap into the rotor side and making this part\u2019s permeability become smaller. Furthermore, different permeability portions will have different eddy current decay speeds. Fig. 2 shows the relative permeability distribution in a dynamic process, due to the existence of a large number of periodic eddy currents, areas with high saturation will focus on tooth top, slots wall and slots bottom. As for the third assumption, in most related literatures, the air-gap synthetic magnetic potential is used to reflect the interaction of d-axis and q-axis currents. However, it introduces the mutual inductance Ldq and Lqd, causing the two axes cannot be decoupled. Meanwhile, Ldq and Lqd need to be obtained by calculating the magnetic energy storage, which not only greatly increases the complexity of the model, but also has low practicability when applied in power system analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003206_s0869864318050104-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003206_s0869864318050104-Figure3-1.png", + "caption": "Fig. 3. Profiles of beads deposited on one another.", + "texts": [ + " Cladding of beads in the form of a wall Consider the particular cases when the shape of the substrate surface is specified analytically and one can integrate equation (7). One of them arises when the surface of already built bead is the base surface. This is a typical case in laser cladding when one forms from a powder, for example, a thin vertical wall [1]. Since the upper ith bead contacts only the lower (i\u22121)th bead, it is sufficient to consider the system of two beads, which is presented in dimensionless coordinates in Fig. 3. The beads are located symmetrically with respect to the OZ axis, therefore, unlike the case depicted in Fig. 2, the left and right corners at the contact points measured along the horizontal coincide: 1 2 ,i i i\u03b8 \u03b8 \u03b8= = i = 1, 2. Because the shape of the lower bead surface is described by analytic equations (3) and (4), one can integrate the balance of forces (7) explicitly and obtain equation (10) in the dimensionless form. Writing the additional integral relation (11) for the bead width ,ib we obtain a system of two nonlinear equa- tions for determining the unknowns 0 , :i iz \u03b8 Fig. 2. Diagram of the single bead profile deposited onto an uneven base with a curved surface shape. Note that relation (10) describes the difference of areas below the upper and lower curves (Fig. 3). To solve the system of equations (10) and (11) it is necessary to specify the angle 1i\u03b3 \u2212 characterizing the side point of the beads contact, which may be found from the equation ( ) ( ) ( )( ) 1 210 0 cos 2 . 2 1 cos i i i d b z \u03b3 \u03be \u03be \u03be \u2212 \u2212 = + \u2212 \u222b (12) 1.4. Cladding of overlapping beads At the laser cladding for obtaining a continuous coating the case occurs frequently when the beads are formed by an overlapping cladding. In such a situation, the beam is shifted to the right by the value d (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002898_s11837-018-3223-3-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002898_s11837-018-3223-3-Figure1-1.png", + "caption": "Fig. 1. CAD rendering of the microscale selective laser sintering system (l-SLS) assembly.", + "texts": [ + " However, these microscale sintering solutions run into problems when the desired feature size becomes less than 10 lm, and nanoscale powders are required to the microscale powders generally used in these systems. For example, particle agglomeration and write speed both become significant problems as the feature size is scaled down below 10 lm. Therefore, this paper presents the subsystems required for a novel microscale additive manufacturing process known as microscale selective laser sintering (l-SLS), which is capable of producing 3D metal parts with sub-10-lm resolutions. Figure 1 shows the designed assembly of the lSLS system which consists of the following subsystems: (1) the slot die coating subsystem to generate a uniform nanoparticle (NP) bed, (2) an optical system to direct and focus laser light to pattern features on to the bed, (3) a laser system to sinter the particles, (4) an XY nanopositioning system to step the bed under the optical system in order to enable large area patterning, and (5) a global travel system to position the substrate between the optical subsystem and the coating subsystem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-26-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-26-1.png", + "caption": "Fig. 19-26: Anti-parallel circuit with circulating current", + "texts": [ + " The motor load is in any case a maximum of the half load of the lift. We assume: a) that the fully loaded motor raises the load by rotating clockwise and we call this \u201cquadrant 1\u201d Determine in which quadrant the lift is operating when: b) the lift is fully loaded when it is lowered c) the lift is empty when lowered d) the lift is empty when raised. SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR 19.23 Fig. 19-25 shows the control circuit of a four quadrant configuration. We refer to this circuit as circulating current free as opposed to the configuration in fig. 19-26 and 19-27. ! 19.24 SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR To ensure a speedy transition between the different quadrants both bridges are operated simultaneously. This is studied in fig. 19-26. At every instant we ensure that \u03b1 1 + \u03b1 2 = 180\u00b0. In quadrant 1 bridge A is operated as rectifier and delivers current I 1 while bridge B is operated as inverter with no current. If bridge A has for example \u03b1 1 = 60\u00b0, then the output voltage is: V 1 = 3 \u0387 \u0302 v ____ \u03c0 \u0387 cos \u03b1 1 = 3 \u0387 ^ v ____ 2 \u0387 \u03c0 . The amplitude \u0302 v here is that of the line voltage of the three-phase power grid. The motor is supplied with voltage V 1 , draws a current I 1 = I a and has a counter emf E 1 with the same polarity as V 1 and in addition E 1 = V 1 \u2212 I a \u0387 R i . If bridge B is operated simultaneously as inverter with \u03b1 2 = 120\u00b0 ( = 180\u00b0 \u2212 \u03b1 1 ) then the average voltage is V 2 = 3 \u0387 ^ v ____ \u03c0 \u0387 cos 120\u00b0 = \u2212 3 \u0387 ^ v _____ 2 \u0387 \u03c0 with the polarity shown in fig. 19-26. Since V 2 is as large and opposed to V 1 , bridge A can only deliver current to the armature of the motor. Bridge B cannot supply current to the motor since the thyristors are polarized in the wrong direction. During braking with the same direction of rotation V 1 < E 1 and bridge A (rectifier) is without current. With E 1 > V 2 bridge B will carry current. With \u03b1 1 = 45\u00b0 (bridge A = rectifier) then \u03b1 2 = 135\u00b0 (B = inverter). Bridge A supplies the machine as a motor. In a 3x400 V net then ^ v = \u221a ___ 2 \u0387 400 = 565V", + " Bridge B cannot supply the motor since the thyristors are blocking. A quick change of \u03b1 1 to 60\u00b0 ( \u03b1 2 = 120\u00b0) gives: V 1 = 269.8V = V 2 . Since the machine does not slow down instantly, E 1 does not change much and as long as E 1 > V 2 bridge B will operate as an inverter and the machine will brake and operate as a generator. E 1 cannot send current through bridge A since the thyristors are blocking. At the instant that E 1 < V 1 bridge A will operate as rectifier and the machine will operate as a motor. The configuration in fig. 19-26 whereby the firing angle \u03b1 1 of bridge A is 60\u00b0 and \u03b1 2 of bridge B is 120\u00b0. The supply voltage is 3x230 V - 50 Hz. Determine the instantaneous output voltages v 1 and v 2 of both bridges at the times corresponding with \u03c9t = 135\u00b0 and \u03c9t = 165\u00b0. The reference voltage is always v 13 , it goes through zero at t = 0. 1. \u03c9.t = 135\u00b0 Fig. 8-21: v 1 = v 13 = 230 \u0387 \u221a __ 2 \u0387 sin 135\u00b0 = 230V Fig. 8-25: v 2 = v 12 = 230 \u0387 \u221a __ 2 \u0387 sin (135\u00b0 + 60\u00b0) = \u2212 84.18V 2. \u03c9.t = 165\u00b0 Fig. 8-21: v 1 = v 13 = 230 \u0387 \u221a __ 2 \u0387 sin 165\u00b0 = 84,18V Fig. 8-25: v 2 = v 12 = 230 \u0387 \u221a __ 2 \u0387 sin (165\u00b0 + 60\u00b0) = \u2212 230V In reality the average output currents of both bridges can be equally large but the instantaneous values of the output voltages are different. This is clearly seen in numeric example 19-3. The result of this is that equalizing or circulating current I k will flow. To limit these circulating current, chokes are used. In fig. 19-26 the current I a + I k will flow through L 1 and L 3 . To avoid saturation large and expensive coils would need to be used. Small coils are used however which saturate for I a + I K but not for I k . The unsaturated coils L 2 and L 4 limit I k . In the third quadrant L 1 and L 3 limit the circulating current while the other coils are saturated. The advantage of speed control with circulating current is that the current in the armature can change direction quickly. The advantage of this is good dynamic control behaviour" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001379_peds.2009.5385887-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001379_peds.2009.5385887-Figure4-1.png", + "caption": "Fig. 4. Magnetic field distributions under different operation modes. (a) DSPM mode. (b) SR mode.", + "texts": [ + " In order to be more applicable for dynamic analysis, the circuit-field-torque coupled time-stepping FEM is usually adopted [28,29]. Moreover, because of frequent magnetization and demagnetization processes of the AlNiCo PMs in the DC-excited memory motor, the Preisach hysteresis model should be incorporated into the time-stepping FEM for analysis [30]. Firstly, by using the FEM, the magnetic field distributions of the proposed DC-excited memory motor under both the DSPM mode and the SR mode are shown in Fig. 4. It can be found that there are 4 phase windings conducting simultaneously at any instant under the DSPM mode, whereas there are 2 phase windings conducting simultaneously at any instant under the SR mode. Further, Fig. 5 shows the self inductance characteristics of the motor under the DSPM mode and the SR mode, which reveals that the self inductance in the SR mode is larger than that in the DSPM mode due to the lack of PM flux. Secondly, the simulated and measured back electromagnetic force (EMF) waveforms of the proposed memory motor at the speed of 200 rpm with respect to the PM flux coefficient \u03ba are both shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002873_smasis2018-7915-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002873_smasis2018-7915-Figure1-1.png", + "caption": "FIGURE 1: SCHEMATIC REPRESENTATION OF (A) CLASSICAL AIR FOIL BEARING COMPARED TO (B) ADAPTIVE AIR FOIL BEARING IN THREE-PAD ARRANGEMENT WITH CORRESPONDING COORDINATE SYSTEM", + "texts": [ + " In order to include the elastic attachment of the bearing shells a stiffness matrix is calculated from a finite element model. The matrix allows to consider the potential energy in the elastic attachment. The model is utilized to perform parametric studies on the bearing shells which are basis for a shape optimization. It can also be used to evaluate the bearing shell\u2019s reaction on the pressure distribution in the air film. 2 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/15/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Fig.1 compares the classical air foil bearing of the bump-type with corresponding adaptive concept, in which an adaptable mechanism is used to actively manipulate the inner contour of the bearing. A classic foil bearing consists of a rigid housing and foil structure (bump-strip and top-foil). In an adaptive air foil bearing, the foil structure is supported by a flexible shell (or segment) and joints between the segment and the housing. On the outside of each supporting shell, there is a piezoelectric patch that generates the required mechanical tension in the circumferential direction of each pad to change the shape (radius and center of curvature) of the bearing clearance on demand" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002713_chicc.2018.8484123-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002713_chicc.2018.8484123-Figure2-1.png", + "caption": "Fig. 2: Joint coordinate of right arm", + "texts": [ + " The number of each robot arm\u2019s DOF is four, but the number of variables required to describe a given task is three, i.e., the x, y, z variables in the operating space for describing the position. Obviously, the latter is less than the former. Therefore, the robot arm is kinematics redundancy. The mathematical model of the dulcimer music-playing robot is established firstly. According to the relative position relationship between joints of the robot arm, the right arm is taken as an example to establish the arm joint coordinate system. Fig. 2 illustrates the graphical view of all joints with joint values (\u03b81, \u03b82, \u03b83, \u03b84). Assume that, \u03b8i is the angle of joint i, \u03b1i is the angle between link i and link i \u2212 1, ai is the length of robots link i, di is the distance between link i and link i\u2212 1, i = 1, 2, 3, 4. The Denavit Hartenberg (D-H) representation is a systematic notation for assigning right-handed orthonormal coordinate frames to each link in a kinematics chain [15, 16]. The modified D-H parameters of the right arm and joint regions shown in Table 1 of the right robot arm can be defined by the joint coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000366_msf.626-627.517-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000366_msf.626-627.517-Figure1-1.png", + "caption": "Fig. 1 (a) Model of the rotor bearing system with rotor/stator rub-impact and oil whirl; (b) Model of the rub-impact", + "texts": [ + " The periodic motion of such rotor bearing are obtained by continuation-shooting method for the periodic solution of nonlinear non-autonomous system, and the bifurcation and stability of the periodic motion are analyzed on the basis of Floquet theory. The effect of the unbalance and rotor/stator clearance on the bifurcation and stability of the periodic motion of the system are investigated. 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: 128.6.218.72, Rutgers University Libraries, New Brunswick, USA-12/04/15,04:47:02) The discussed model shown as Fig. 1 is a simple Jeffcott rotor supported on two equal cylinder short journal bearings, and consists of a massless elastic shaft carrying a disk at the middle of the span. O1, O2, O3 are the geometric center of the journal bearing, the geometric center of the disc and the mass center of the disc respectively. m1 is the equivalent mass of the journal bearing , and m2 is the equivalent mass of the rotor disk. k is the stiffness of the rotor shaft. c1 is the damping at the bearing , and c2 is the damping at the rotor disc. Rub Impact Forces. It is assumed that the time during rub impact is very short compared with one complete period of rotating; therefore, an elastic impact model is used. Also the Coulomb type of frictional relationship is assumed in the analysis. When rub happens as shown in Fig. 1(b), radial impact force F and tangential rub force TF can thus be expressed as ( ) ( ) 0 (for ) , (for ) c e F x y e k e \u03b4 \u03b4 \u03b4 <= \u2212 \u2265 T F fF= (1) where 2 2e x y= + , ck is the stiffness of the stator, \u03b4 is the initial clearance between rotor and stator, and f is friction coefficient of rotor and stator. These two forces can be written in x , y coordinates c 1( - ) ( ) 1 0 ( ) x y x y P f xe k e P f ye P P e \u03b4 \u03b4 \u03b4 \u2212 = \u2212 \u2265 = = < (2) Equation indicates that when e is smaller than \u03b4 , there will be no rub impact interaction and the rub impact forces are zero while the rub impacting will happen if e is bigger than \u03b4 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002643_chicc.2018.8483091-Figure19-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002643_chicc.2018.8483091-Figure19-1.png", + "caption": "Fig. 19. Generic Nascap-2k geometrical model used to estimate charging on Galaxy 15.", + "texts": [ + " Also, housekeeping functions were unimpaired, so it maintained power and operations until its momentum wheels finally saturated in December, 2010, and it underwent a reset and became operational once more. Although the cause of the anomaly remains officially proprietary, it was most likely caused by the space environment because of the timing of the failure. In what follows, we recount what was known by the authors as of January, 2011 [19]. To estimate the charging that occurred on Galaxy 15 in a nonproprietary manner, a generic Nascap-2k geometrical model was used that incorporates many features of commercial GEO satellites. It is shown in Fig. 19. In the Nascap-2k modeling, the uneclipsed Sun was allowed to impinge on the solar cells, and in eclipse the Sun was turned off. Material properties were the default values for Nascap-2k. The satellite environment history was kindly furnished by J. Rodriguez as measured by the NOAA GOES-13 and GOES-14 satellites. We assume that Galaxy 15, only about 30\u25e6 in longitude from the nearest GOES satellite, experienced the same electron and ion environment. In the case of Galaxy 15, the geomagnetic storm impacted on the magnetosphere while the satellite was in eclipse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000230_20080706-5-kr-1001.01206-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000230_20080706-5-kr-1001.01206-Figure2-1.png", + "caption": "Fig. 2. Principle of camber direct measurement", + "texts": [ + " This is in order to tolerate wheel cornering and damper deflections. When the sensor unit is fixed on the mechanical adapter, its displacements with respect to the road (and to car body) are the same as those of the wheel rim, despite the tire solicitations. Thus, the measured angle is the angle between the rim plan and the road, called wheel/road angle (\u03b3wheel/road). Once this angle is measured, the camber angle (\u03b3) can be easily determined with the following relation (1): \u03b3 = \u03b3wheel/road \u2212 90\u25e6 . (1) Fig. 2 shows the principle of camber direct measurement via 3 distance sensors. Many industrial sensors use this measurement principle. They differ by the precision of distance measurement and two distance sensors only can be used. An application example is presented in Nu\u0308ssle and Gnadler [2001]. A prototype sensor using 3 distance sensors was developed in the laboratory and results are presented in Basset et al. [2005]. L1 2 = (L1 \u2212 L2) \u2217 ( a + b a ) , \u03b3wheel/road = arctan ( L3 \u2212 L1 2 c ) . (2) Indirect measurement This method requires two different measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000592_cdc.2008.4739453-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000592_cdc.2008.4739453-Figure2-1.png", + "caption": "Fig. 2. Input-output behavior for the actuator with saturation and hysteresis", + "texts": [ + " (5) In the following subsection, we will describe the behavior of the system with hysteresis, which can be regarded as a special case of the aforementioned hybrid model. Consider the input-output behavior of an actuator v = u \u2212 \u01ebq, \u2200(u, q) \u2208 {(u, q) \u2208 R \u00d7Q\u2032 : uq \u2264 a}, (6) where u is the feedback control input, \u01eb > 0 is the width of the hysteresis, a > 0 is the threshold and q \u2208 Q is the logic variable. Fig. 1 shows the input-output relationship. On the other hand, when the actuator is saturated, its inputoutput behavior is shown in Fig. 2 and the representation is given as v = sat(u \u2212 \u01ebq), \u2200(u, q) \u2208 {(u, q) \u2208 R \u00d7Q\u2032 : uq \u2264 a}, (7) where \u01eb and a are defined in the same way as below (6). Assumption 1: a > 0, \u01eb > 0, \u01eb < min{1, a} and 1 \u2212 \u01eb < a \u2264 1 + \u01eb. Hence, in Fig. 2, the threshold is such that switching happens in the linear section of each curve, and after switching, the output is saturated. When a = \u01eb+1, two curves connect and the output does not jump. To model the input-output behavior of the actuator in a control system, we define u\u0303 = u \u2212 \u01ebq such that \u03c3 = v. By appropriately choosing the control framework and defining the coordinates, we are able to embed the system with hysteresis into H = (f, C, g,D) description and apply statefeedback laws. Later in Section IV, the modeling procedure will be illustrated by using a numerical example for a tracking problem", + " The control objective is to achieve exact tracking of constant references r = 2/3 and the closed-loop poles are all located at s = \u22121. To implement the integral control for exact tracking (see e.g., [2, page 552]), denote e = x1\u2212r and xI = \u222b edt. Then y = e + r and the system is e\u0307 x\u03072 x\u0307I = 0 1 0 0 \u22120.5 0 1 0 0 e x2 xI + 0 1 0 u. (17) The control law is given by u = \u22123e \u2212 2.5x2 \u2212 xI so as to stabilize the system (17) and achieve exact tracking for the original system. However, if such a system is subject to actuator nonlinearity as shown in Fig. 1, and Fig. 2, then the closed-loop performance will be affected. As in Section II, we describe this system with a hybrid model as follows. States: [\u03be1 := e, \u03be2 := x2, \u03be3 := xI + \u01ebq, q ] T . The flow map is given as: f(\u03be, q) = [\u03be2, \u22120.5\u03be2 + \u03c3(u\u0303), \u03be1, 0] T , where u\u0303 = u\u2212\u01ebq = [ \u22123 \u22122.5 \u22121 ] \u03be. In section IV-A, we will discuss the case \u03c3(u\u0303) = u\u0303, while in section IV-B, we will discuss the case \u03c3(u\u0303) = sat(u\u0303). The jump map is: g(\u03be, q) = [\u03be1, \u03be2, \u03be3 \u2212 2\u01ebq, \u2212q] T . The flow set and the jump set are: C := { [ \u03be q ] : [\u22123 \u2212 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000226_j.1747-1567.1988.tb02171.x-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000226_j.1747-1567.1988.tb02171.x-Figure1-1.png", + "caption": "Fig. 1 --Bearing outer ring deflection illustrated", + "texts": [ + " A standard loaddeflection curve for a ball bearing gives the relation between applied force and total relative motion between the inner ring/shaft and outer ring/housing. This is the \u2018macro\u2019 deflection curve. For thrust-loaded bearings, the macrodeflection is measured in the axial direction. Ball bearing rings are of relatively thin cross-section compared to the contact stresses they support. Contactstress levels can commonly be expected to be greater than 100,000 psi and often in excess of 200,000 psi. Forces transferred through the ball/raceway contacts cause bending and localized deflections on the bearing rings as illustrated in Fig. 1. The deflections are centered over each ball, varying in amplitude across the width of the bearing, and \u2018travel\u2019 about the ring with the ball set during bearing operation. \u2018MICRO\u2019 DEFLECTION The \u2018micro\u2019 deflection of the bearing is that localized radial deformation of the bearing ring which occurs at each ball location with the application load. Strain gages can be bonded to bearing rings and used to measure flexure of the ring due to loading, however the bearing housing and the bearing itself must be modified to accommodate the gages and their wires" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000089_j.jappmathmech.2009.08.012-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000089_j.jappmathmech.2009.08.012-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " The phenomenon of the rotation frequency of the actuating system \u201cbecoming stuck\u201d at a constant value in the vicinity of the resonance frequency of the vibrating system, which has been named the Sommerfeld effect,4 also occurs in the case of rotating shafts. This effect also explains the build up of the vibrations of a shaft in the resonance parameter zone. The difficulties in an analytical investigation of \u201cflexible-rotor/actuating-mechanism\u201d dynamical systems are attributed not only to their non-linearity and cylindrical phase space, but also to their high dimension. In particular, a system consisting of an unbalanced shaft interacting with a limited power source has three degrees of freedom in the single-mode approximation (Fig. 1). In such a case the method of averaging is actually the only analytical method for investigating the system dynamics. However, its use in the standard implementation (the use of amplitude\u2013phase replacements, which bring the system into a standard form) leads to averaged systems whose investigation is a no less a problem than the investigation of the original systems.5,10 The method is effective in the proposed form, in which the dimension of the system is reduced to half, and the averaged system is specified in three-dimensional phase space", + "5,10,17 Nevertheless, because of the importance of this model, particularly from the practical standpoint, we will return to it and investigate its dynamical properties in detail. We will assume that the mass distributed along the shaft length is reduced to the mass of a disk located at the middle of a weightless shaft that is not torsionally compliant. The disk eccentricity is denoted by e, W is the geometric centre of the disk, S is the centre of gravity, and O is the equilibrium position of the disk (Fig. 1). Prikl. Mat. Mekh. Vol. 73, No. 4, pp. 552-561, 2009. \u2217 Corresponding author. E-mail addresses: s.verichev@mail.ru (N.N. Verichev), n.verichev@yandex.ru (S.N. Verichev), erf04@sinn.ru (V.I. Yerofeyev). 0021-8928/$ \u2013 see front matter \u00a9 2009 Elsevier Ltd. All rights reserved. doi:10.1016/j.jappmathmech.2009.08.012 In dimensional variables and parameters the system dynamics is described by the equations5,10 (1.1) Here x and y are the coordinates of the centre of gravity of the disk in the fixed system of coordinates with origin on the axis of the unperturbed shaft and the (x, y) plane perpendicular to the axis, and k are the external and internal damping factors for flexural vibrations, c is the shaft stiffness at the point of attachment of the disk, m is the mass of the disk, J is the moment of inertia of the motor rotor, is the range of rotation, and L(\u0307) is the torque of the motor, which includes the moment of the internal forces of resistance to the rotor motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003662_978-3-030-13273-6_43-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003662_978-3-030-13273-6_43-Figure5-1.png", + "caption": "Fig. 5. Diagram of the 3D scanning process with the use of the GOM system [16].", + "texts": [ + " The reference model is a small body made of steel with main dimensions: 70 L 50 W 25 H mm, built of freeform surfaces (Fig. 4). The model has been scanned twice. The first scan was captured with the ATOS Core 80 scanner (Fig. 6a) and was adopted as a reference measurement, performed with the most accurate of currently available methods on the market. All accuracy comparisons in this paper were made with respect to this measurement. The second scan was acquired with the ATOS III Triple Scan device (Fig. 6b) and was used to examine the accuracy of structured-light scanning (Fig. 5) in the reverse engineering process and to generate the digital model for the subsequent manufacturing process. Both scanners use the ATOS Professional software for the acquisition and preprocessing of data. Exposure time is an essential parameter that can be modified during scanning and which has a direct impact on the time and the accuracy of the scanning process. For the software employed, it is possible to perform scans with one, two or three exposure times. A single exposure time is designed for simply shaped objects that reflect light well" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003460_s1052618818060110-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003460_s1052618818060110-Figure1-1.png", + "caption": "Fig. 1. Scheme of the ultrasonic burnishing through the example of a rod: 1, workpiece; 2, indenter; 3, wave conductor; 4, cartridge; 5, ultrasonic transducer; VR is the rotational speed of the work piece; FN is the pressing force of the indenter; \u03bem is the amplitude of the indenter displacement.", + "texts": [ + "5, at room temperature at a rate of 4 m/min in the step mode with adjustable single gage reduction (~25 \u03bcm). After each rolling step, the samples were cooled down in water in order to avoid the inf luence of the deformational heating. The cumulative value of the true strain was \u0435 = 1.9 (\u0435 = ln t0/tf), where t0 and tf are the initial and final thickness of the plate. The after strain annealing was carried out at 450\u00b0\u0421, and the holding time was 1 h. The thickness of samples after rolling was 0.3 mm. Then the samples were subjected to ultrasonic burnishing1; the scheme is given in Fig. 1 [11]. The ultrasonic burnishing was carried out with the following parameters: a fixed frequency of 20 kHz; an amplitude of free oscillations of the tool, \u03bem = 20 \u03bcm and a diameter of the indenter ball of 10 mm, a processing rate of VR = 0.3 m/min, the steady-state force of the pressure of the tool to the work surface FN = 50, 100, and 150 N, and a tool advance of 0.1 mm. Samples before and after burnishing are shown in Fig. 2. The study of the microstructure was made with an optical microscope and JEM-2000 transmission electron microscope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002786_embc.2018.8512702-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002786_embc.2018.8512702-Figure4-1.png", + "caption": "Fig. 4. Operated object equipped with a force sensor", + "texts": [ + " Note that, we exported the motion data to BVH format which is a commonly used format of motion data. Other motion capture device can also be used for our method if it can export the motion data as this format, such IMU based motion suit, Kinect, and so on. We calculate the positions of each joint and the CoM from the motion data (BVH format) by using a wildly used human dynamic model - OpenSim. For measuring the real operating force, we equipped a 6-axis force-torque sensor (DynPick WDF-6A200-4-AG2, WACOH) on the box (Fig. 4). The dumbbells were put on the box for changing the weight. When the subjects push or pull the handle, the operating force was measured by the force sensor. For the lifting experiment, we rotated the box and make the handle upside. We put the dumbbells in the box. The subject held the handle and lifted the object up. The experiment was conducted on three male subjects. Every motion (pushing, pulling and lifting) have been carried out for three different weights. Each weight had been tested five times" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002483_aim.2018.8452437-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002483_aim.2018.8452437-Figure7-1.png", + "caption": "Fig. 7: Schematic view of experimental setting for plate bulldozing test.", + "texts": [ + " From a metal plate bulldozing experiment and simulation results, validation of Hegedus\u2019s model is confirmed in our experiment environment. From a wheel bulldozing experiment and simulation results, the applicability of Hegedus\u2019s model for a wheel and relationship between a bearing force and a sinkage is confirmed. In order to confirm validation of Hegedus\u2019s model, a metal plate bulldozing tests conduct. 1) Experimental environment and conditions: The overview of the experimental system and environment are shown in Fig. 7, 8. The bulldozing area has a width, length and height of 300 [mm], 1200 [mm] and 0\u2212 30 [mm], respectively, and is filled with Silica sand No. 5 as loose soil [22]. The bulldozing force is obtained by force sensor, which sets up upper of the metal plate. The rope, which connects to the pole, pulls the metal plate. The soil specific parameters are listed in Table II. The metal plate size is 100 [mm] width, 8 [mm] length and 70 [mm] height. The bulldozing speed sets at 36.5 [mm/s] and sinkage conditions set at 0\u221230 [mm]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001030_icit.2008.4608632-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001030_icit.2008.4608632-Figure1-1.png", + "caption": "Fig. 1. Structural of dual-three-phase induction machine", + "texts": [ + " MACHINE MODEL To model the dual three-phase induction machine, we assume that the following assumption is tenable: i) The rotor cage is equivalent to a six-phase wound rotor and the windings are sinusoidally distributed. ii) The magnetic saturation and the core losses are neglected and the mutual leakage inductances of different phase windings in the same slots are neglected. iii) Flux path is linear. Assume neutrals of two sets of three phase stator winding are connected, the 1st and 6th stator phase windings are open circuited as shown in Fig. 1. In this case, we define the stator voltage vector [vs] and the current vector [is] as follows: [ ] [ ] [ ] [ ] 2 3 4 5 2 3 4 5 T s s s s s T s s s s s v v v v v i i i i i = = Since the rotor remains as its original structure, the [vr ]vector and the [ir ]vector keeps unchanged as in balance as follows: [ ] [ ] [ ] [ ] 1 2 3 4 5 6 1 2 3 4 5 6 T r r r r r r r T r r r r r r r v v v v v v v i i i i i i i = = The stator resistance matrix [Rs] and rotor resistance matrix [Rr] define as follows: [ ] [ ] [ ] [ ]4 4 6 6 s s r rR r I R r I \u00d7 \u00d7 = = Where [I] is an identity matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000806_icis.2009.184-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000806_icis.2009.184-Figure3-1.png", + "caption": "Figure 3. Membership functions for yaw rate ( )", + "texts": [], + "surrounding_texts": [ + "Keywords-intelligent control; fuzzy control; neurofuzzy; ship autopilot\nI. INTRODUCTION Since last two decades, studies on automatic ship berthing have been carried out by many researchers [1]-[5]. This topic of study is one of the difficult problems in ship control fields[1]. Therefore, almost recent researches in automatic berthing control tried to employ \"intelligent control\" that can in some extents mimics human operators. These control techniques include knowledge-based control systems, expert systems, fuzzy logic controllers and neural network-based controllers. It has been proved that, neural network (NN) is an effective and attractive option in developing automatic ship berthing controllers.\nThe classical control theory was applied to steering control. The first automatic ship steering control system was reported. It was based on a simple proportional controller which to ensure stable operation under all conditions, employed low gains with inferior accuracy of course keeping. When the PID controllers became commercially available, they greatly improved the performance and until the 1980s all makes of autopilot were based on these controllers. The main disadvantage of the PID controllers is that they require manual adjustment to compensate for changes of operating conditions, but these settings are usually not optimal for the ship. The adjustments are time consuming and tedious. Such autopilots also perform badly in rough seas. Hence there was a need to design autopilots that could automatically adjust or adapt themselves. Adaptive control not only makes the ship operation more convenient but also minimizes the drag due to steering. Many adaptive algorithms for a ship steering control are therefore available in the literature. In the recent years, sophisticated ship autopilots have been proposed based on\nadvanced control engineering concepts. These include model reference adaptive control, self-tuning control, optimal control, neural control and fuzzy control [6-9].\nThe main difficulty in discussing intelligent approaches to autopilot design is that of defining what it is meant by intelligent or intelligent control and stems from the fact that there is no general agreement upon definitions for human intelligence and intelligent behaviour. One of the earliest definitions of the criterion for machine intelligence, is that by Turing who expressed this criterion as the well-known Turing test. The Turing test is undertaken in the following manner: A person communicates with a computer through a terminal. When the person is unable to decide whether she/he is talking to a computer or another person, the computer can safely be said to possess all the important characteristics of intelligence. The Turing test clearly emphasises the external behaviours requested for a computer (or machine) to be safely defined as an intelligent machine. These external behaviours do not have to be distinguished from that of a human being and most importantly this has to happen from a human point of view (that is the observer). The problem of differentiating between an external manifestation of such behaviours and the internal mechanism that produces it, is also mentioned by Zadeh as the main reason why a clear definition of adaptation is still lacking. The viewpoint adopted throughout this paper is that intelligent control is the discipline that involves both artificial intelligence and control theory. The design of intelligent control systems should be based on an attempt to understand and duplicate some or part of the phenomena that ultimately produces a kind of behavior that can be termed intelligent, i.e. generalisation, flexibility adaptation etc. The two paradigms which are generally accepted as matching or replicating these intelligent characteristics are fuzzy logic and artificial neural networks. Over the last twenty years or so there has been an explosion in interest in applying intelligent control to a wide range of application areas, including ship autopilot design. The motivation for the work on developing intelligent ship autopilots has been a wish to utilise the powerful approximation and learning abilities of such approaches to produce intelligent autopilots which in terms of their robustness qualities can more effectively cope with the non-linear and uncertain characteristics of ship steering.\n978-0-7695-3641-5/09 $25.00 \u00a9 2009 IEEE\nDOI 10.1109/ICIS.2009.184\n701", + "A summary of some of the approaches is presented in the following sections.[10]\nII. FUZZY CONTROL APPROACHES FOR SHIP AUTOPILOT The use of fuzzy set theory as a method for replicating the non-linear behavior of an experienced helmsman is perhaps the most appropriate application of this technique. Fuzzy rules of the type: F heading error is positive small AND heading error rate is positive big THEN rudder angle is positive medium, typify the actions of an experienced helmsman. The schematic of a fuzzy logic controller is shown in Fig. 1 where the conventional controller block is replaced by a composite block comprising four components:\nThe goal of the designer is to ensure that the output matches the reference. The memberships have been defined in the two inputs and one output fuzzy system. The inputs and outputs to the fuzzy controller are The yaw a error( ) ( reference ship course subtracted from actual course.) The yaw rate ( ) (previous error subtracted from current error) over one sample period. The output of the controller is the rudder angle ( C ).\nIn practical applications, the control goals and system constraints are all of fuzzy characters, in order to unify them, fuzzy membership function is used to express their characters. These operators can be used to translate a linguistic description of control goals into a decision function. In this way, various forms of aggregation can be chosen giving greater flexibility for expressing the control goals. The universe of discourse (range) of the inputs and outputs are mapped into several fuzzy sets of desired shapes. The membership functions for the inputs are shown in Fig 2,3 and outputs are shown in Fig 4.\nA fuzzy system is characterized by a set of linguistic statements based on expert knowledge. The expert knowledge is usually as \u201cif-then\u201d rules, which are easily implemented by fuzzy conditional statements in fuzzy logic. Fuzzy control rules have the form of fuzzy conditional statements that relate the state variables in the antecedent and process control variables in the consequence. Rules that were developed in the work are given in Table1.\nThe output is defuzzified to get a final value of the control. The widely used the center of area method strategy generates the center of gravity of the possibility distribution of a control action. the control surface is shown in Fig 5.\nFor course-changing and course-keeping of the ship manipulating system, the performance of the fuzzy controller designed is implemented in MATLAB/Simulink environment.", + "For large signal transient in course-changing, the response to course-change manoeuvres of 20 degrees were obtained. The yaw response of the fuzzy controller was shown in Figure 6. For course-keeping manoeuvres in the presence of external disturbances, the yaw response of the fuzzy controller designed under the worst case conditions of the ship system which exist random disturbance was shown in Figure 7.\nOne of the main advantages of fuzzy logic autopilots is that the rules may be formulated without a precise definition of the ships dynamics. The control actions are normally decided such that the required rudder demand is determined from knowledge of the expected ships response to the input. Although this process is best achieved through consultation with the helmsman or by observation of his/her actions, the fuzzy rules may be formulated from an understanding of the systems dynamical behaviour. This process illustrates the fundamental difference of fuzzy controller design compared with more traditional model-based controller designs. Whereas the latter is concerned with designs to meet performance specifications, i.e. damping, speed of response, steady-state error, etc., fuzzy controller design is focused on predicting system behaviour in response to specific inputs, (as postulated by Sperry and Minorski following their observations of the helmsman). What this means is that often fuzzy approaches are application specific and the final fuzzy controller is often arrived at after considerable trial and error where the distribution and shape of the fuzzy sets used are tuned in order to achieve the desired performance. However it should be noted that fuzzy controllers are intrinsically robust, in the sense that through the generalisation property they can accommodate new\nsituations (i.e. small changes in system parameters). They are also non-linear, in that they incorporate (map) the functional relationship between input and output which is a non-linear relationship.\nUsing a neural network to mimic the operation of a fuzzy autopilot is an option, but this direct copying of the fuzzy rules by a neural network results in a controller which looses the transparency of the original, which is replaced by a black-box network. The attractiveness of combining the transparent linguistic reasoning qualities of fuzzy logic with the learning abilities of neural networks to create intelligent self-learning controllers has, over the last decade, led to a wide range of applications to be reported in the literature. Such approaches have brought together the inherently robust and non-linear nature of fuzzy control with powerful learning methods through which the deficiencies of traditional fuzzy logic designs may be overcome. Many of the proposed fusions may be placed into one of two classes: either networks trained by gradient descent or reinforcement paradigms, although some methods combine these learning techniques. Whatever training method is chosen the parameters within the fuzzy controller which are to be tuned must be selected. One of the most useful and much used combinations of neural networks and fuzzy logic is the Adaptive Network Based Fuzzy Inference System (ANFIS) proposed by Jang whereby, d, the fuzzy consequences of first-order Sugenotype rules of the form: If are tuned using neural networks.\nANFIS performs static non-linear mapping from input to output space but without modification it cannot be used to represent dynamic systems. In order to identify dynamic systems, a combination of ANFIS with some time delay units and feedback is required. Hence, non-linear dynamic system can be modeled by ANFIS combined with some time delay units. An excessive number of inputs not only impair the transparency of the underlying model, but also increase the complexity of computation necessary for building the model. Therefore, it is necessary to do input selection that finds the priority of each candidate inputs and uses them accordingly.\nAnfis is much more complex than the fuzzy inference systems discussed so far, and is not available for all of the fuzzy inference system options. Specifically, anfis only supports Sugeno-type systems. All output membership functions must be the same type and either be linear or constant. Different rules cannot share the same output membership function, namely the number of output membership functions must be equal to the number of rules.\nUsing a batch learning scheme and a hybrid learning rule, i.e. BP algorithm is applied to the learning of premise parameters, while least square algorithm to the learning of consequent parameters, an ANFIS system for ship autopilot" + ] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.12-1.png", + "caption": "FIGURE 16.12 Schematic of conduction mode welding.", + "texts": [ + " The laser power is first deposited on the surface and then transferred by conduction to the surroundings. The penetration is thus essentially controlled by conduction from the initial point of contact, that is, the surface of the workpiece, causing a small surface area to be heated above the melting point. Convection also plays a role once a weld pool is formed. The resulting weld is shallower with a wider heat-affected zone compared to that produced by keyhole welding. It is more common with low-power lasers, say below 1 kW. A typical configuration is shown in Fig. 16.12. The shape of the weld pool in conduction mode welding is influenced by flow in the weld pool and by the presence of surface active elements (see Section 10.3.3). Approximate analysis of heat flow in conduction mode welding can be done using the point heat source analysis discussed in Section 10.2, while material in Section 10.3 will enable flow in the molten pool to be analyzed. At very high power densities, say, above 106 W/cm2, part of the work material is vaporized to form a cavity, the keyhole, which is surrounded by molten metal, which in turn is surrounded by the solid material (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure18-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure18-1.png", + "caption": "FIG. 18: UNROLLED GEOMETRY FOR 3 CAM LOBES.", + "texts": [ + " Offsetting the upper bound by the desired slicing layer height, and trimming these offset curves to the unrolled geometry edges will generate the stop-start points for the deposition segments for each deposition layer. Using a crude offset distance, this is illustrated for the cam profile. Attention must be paid to managing the unrolled geometry initial positions, as this links the angular relationships to the deposition start-stop points for a given roll diameter. The unrolled geometry for a 3-lobed cam is shown in Fig. 18. 8 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use This procedure would work for bosses and recessed features appended on a rotary base via the creation of unrolled / re-rolled edge geometry from axial slices. Appended features more aligned axially need another tool path solution approach. For the spiral geometry shown in Fig. 19, an unroll solution approach would not be ideal as the numerous linear tool paths would have many stops and starts, and rapid moves" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.35-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.35-1.png", + "caption": "Fig. 11.35 Self-aligning torque in a combined slip. Reproduced from Ref. [2] with the permission of Guranpuri-Shuppan", + "texts": [], + "surrounding_texts": [ + "Fx \u00bc Cxbl2h 2 s\u00fe n\u00fe 1 n 2nFzld ln\u00fe 1 l 2 n l lh\u00f0 \u00de 1 n\u00fe 1 l 2\nn\u00fe 1\nlh l 2\nn\u00fe 1 ( )\" #\ncos h\nFy \u00bc bCyl2h s1 2 tan a d Cy \u00fe l2 3r2Ky Fy s2 l 1 2 s2lh 3l\n\u00fe n\u00fe 1 n 2nFzld ln\u00fe 1 l 2 n l lh\u00f0 \u00de 1 n\u00fe 1 l 2\nn\u00fe 1\nlh l 2\nn\u00fe 1 ( )\" #\nsin h\nMz \u00bc FxFy\nK 0 y\n\u00fe bCys1l2h lh 3 l 4 tan a\nbCy d Cy \u00fe l2 3r2Ky Fys2 l2h l lh 3 s2l2h 4l l 4 \u00fe s2lh 6\n\u00fe n\u00fe 1 n 2nFzld ln\u00fe 1 l 2\nn\u00fe 1\nl lh\u00f0 \u00de\u00fe l 2\nn l2 l2h\n2 1 n\u00fe 2 l 2\nn\u00fe 2\nlh l 2\nn\u00fe 2 ( )\" #\nsin h\n\u00f011:130\u00de where sh is determined according to\nn\u00fe 1 n 2nFzls ln\u00fe 1b\nl 2\nn\nlh l 2\nn\n\u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Cxslh\u00f0 \u00de2 \u00feC2\ny lhs1 tan a d Cy \u00fe l2 3r2Ky Fys2 lh l 1 s2 lh l\n2 s\n:\n\u00f011:131\u00de\nNote that lh = 0 is used in the case that lh < 0. The effective slip angle a is smaller than the slip angle of the wheel a0 owing to the torsional deformation of the tire caused due to the self-aligning torque. The effective slip angle a is given by Eq. (11.68).\n(2) Example calculations\nThe computational flow when calculating the side force, fore\u2013aft force and self-aligning torque is the same as that in Sect. 11.2. Sakai [2] calculated the side force and self-aligning torque with lateral and fore\u2013aft forces as shown in Figs. 11.34 and 11.35. The parameters used in the calculation are the contact length (l = 100 mm), contact width (b = 120 mm), radius of the tire (r = 270 mm), shear spring rate of the tread (Cx = Cy = 100 MPa/m), inflation pressure (Fz = 3 kN), velocity (V = 60 km/h), fundamental lateral spring rate of the sidewall (ks = 158 kN/m2), lateral spring rate of the tire (Ky = 50 kN/m), bending rigidity (EIz = 2.5 kNm2), torsional spring rate of the tire (Rmz = 10 kNm/rad), friction coefficient (ls = 1.1, ld = 0.8), friction decreasing constant (aV = 0.005) and inflation pressure (200 kPa) related to the lateral fundamental spring rate of the carcass.\nThe drawback of the friction ellipse (i.e., that the ellipse cannot express cornering properties at a large slip ratio) can be solved using Sakai\u2019s theory.", + "However, Sakai\u2019s theory for a large slip angle and large slip ratio still predicts that the cornering force in the braking condition is stronger than that in the driving condition. Measurements presented in Fig. 11.29 show the opposite relation. This may be because Sakai\u2019s model assumes that the contact pressure distribution does not change even under the fore\u2013aft forces.", + "Miyashita and Kabe [9\u201311] modified Sakai\u2019s theory from Sect. 11.2 by considering the change in pressure distribution under external forces in the circumferential direction. The parameters in the theory are identified by curve fitting the measured load dependency of the force and moment for a small slip angle. Using these parameters, the cornering properties of a tire can be calculated up to a large slip angle and slip ratio.\n11.5.1 Neo-Fiala Model for a Small Slip Angle\n(1) Fundamental equations\nDeformation of the tire tread with a small slip angle in the contact area consists of shear, bending of the belt and torsion of the belt as shown in Fig. 11.36.\nIn the case of a small slip angle, substituting lh \u2245 l into Eq. (11.51), using the second equation of Eq. (11.42), the side force Fy and self-aligning torque Mz are given by\nFy \u00bc Cyl2b 2 tan a dFy\n3lCy\n\u00bc Cyl2b 2 tan a k3lFy\n6ks\n; \u00f011:132\u00de" + ] + }, + { + "image_filename": "designv11_92_0000072_s0036024409080263-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000072_s0036024409080263-Figure3-1.png", + "caption": "Fig. 3. Mechanism of the decomposition of \u00e72\u00e92 on the catalase biomimetic electrode.", + "texts": [ + " The electrodes prepared by the electrochemical method have a limiting sensitivity to the concentration of \u00e72\u00e92, 10\u20132 wt %, while the electrodes obtained by the chemical procedure showed activity in the vibrational mode probably because of diffusion (their sensitivity was 10\u20136 wt %, H2O2). Note that the electrodes prepared by the chemical technique proved rather sensitive, but they were much less stable than the electrodes obtained by the electrochemical method. The supposed electrochemical mechanism of the catalase reaction was stepwise (Fig. 3), as noted in [7]. At the first stage, highly active hydroperoxide particles formed, which were responsible for the transfer of electrons from the cathode to the iron ion and their subsequent distribution. As a result, at the last stage, hydroxide anions were generated to the volume (reaction medium) with concurrent regeneration of the biomimetic. This mechanism shows how pH in the electrochemical system becomes higher than in bidistilled water (pH 6.2). The above-described biomimetic catalase electrodes were studied to examine the peroxidase activity in the transformation of ethanol to acetaldehyde" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000700_cgiv.2009.70-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000700_cgiv.2009.70-Figure7-1.png", + "caption": "Figure 7. Full evaluation of an ant model(298 patches) to depth 5 at 14fps; (a) the original subdivision surface; (b) and (c) deformed surfaces.", + "texts": [ + " By using the OpenGL Interoperability of CUDA, we can process these vertex buffers in GPU computation. After processing, it is directly rendered on GPU. The rendering process is highly accelerated by using vertex buffers. In order to measure the performance of our program, we wobble the vertices of the input mesh along their normals and reevaluate the mesh to the prescribed depth, i.e. 5 in our examples. Figure 1 shows reevaluating a toy model with 66 patches to depth 5 at 56fps. The ant model in Figure 7 with 298 patches is reevaluated to depth 5 at 14fps. More examples are shown in Figure 6(a) - Figure 6(c) on a chessman model with 314 patches and Figure 6(d)-Figure 6(f) on a rocker arm model with 354 patches. All these examples show that our patch-based tessellation algorithm achieves near realtime performance. Compared to our implementation on CPU, the GPU implementation runs about 20 times faster. For instance, the performance for the rocker arm model on CPU is less than 1fps. In this paper, a patch-based tessellation algorithm for DooSabin subdivision scheme is developed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-14-1.png", + "caption": "Fig. 19-14: Single quadrant DC-motor at n = 500 rpm, M t = 0.5 x M = 8.4 Nm and I a max = 9A", + "texts": [ + " Due to the inertia of the motor with its drive in the first instant E remains practically unchanged at E = 52.5V so that: I a = V a \u2212 E ______ R i = 220 \u2212 52.5 ________ 1 = 167.5A ! Too much for our 10 A motor and also for the associated SCR bridge . We are compelled to limit the armature current and maintain it within safe operating values. The only way to do this is to measure the current I a with a current sensor and to control this current to below an adjustable (acceptable) maximum. For this reason the internal current loop of fig. 19-11 is now active. This is shown in fig. 19-14 with the current limited to 9 A, and speed set-point n*= 500 rpm with the motor 50 % loaded. The current sensor is followed by an I/V converter that for example converts 0-10 A to 0-10 V. E = 500 \u0387 e N = 500 \u0387 0.21 = 105V; V a = E + I a \u0387 R i = 105 + 5 \u0387 1 = 110V V di\u03b1 = 110 = 220 \u0387 cos\u03b1 \u2192\u2192 \u03b1 = 60\u00b0 \u2192\u2192 V control = 3.33V SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR 19.17 If the desired speed is 750 rpm then the potmeter wiper is set to7.5V. In the first instant the motor remains at 500 rpm and the error voltage \u03b5 is 2", + " Indeed , if I a > 9A then the error \u03b5 2 would be negative, the PI controller integrates down, V control decreases, V a and I a decrease. With I a = 9A the driving torque of the motor is: M = 90% \u0387 16.8 = 15.12 Nm. M = M t + M v becomes: 15.12 = 8.4 + M v . There is an acceleration torque M v = 6.72 = J m \u0387 d\u03c9 ___ dt . The acceleration d\u03c9/dt depends upon the value of J m . If the motor exceeds 750 rpm then \u03b5 1 is slightly negative and the output of the n-controller drops and after a few short fluctuations it sets the output to 5V and the armature current I a is then 5A (which corresponds to 50% load) and the values of fig. 19-14 are then: . n = 750 rpm ; I a = 5A . \u03b5 2 = 0 . E = 750 \u0387 e N = 750 \u0387 0.21 = 157.5V . output armature current converter = 5V . V di\u03b1 = V a = E + I a \u0387 R i = 162.5V . output current limiter = output n-controller = 5V . \u03b1 = 42\u00b0 . output tacho = 7.5V.. V control = 5.3V Photo Maxon Motor Benelux: Ceramic is better than steel. Possible applications are where steel reaches the limit of its properties. Properties such as resistance to abrasion and good conduction prove that the lifespan are dramatically extended" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000664_icma.2009.5246479-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000664_icma.2009.5246479-Figure6-1.png", + "caption": "Fig. 6. Relationship between the homogeneous matrices and the coordinate frames", + "texts": [ + " The relationship between hand frame {H} and the end effecter frame {E} can be described by rotation matrix : R (8) (9) (II) o o -I while homogeneous matrix ~ R , representing the relationship between end-effecter frame and the master device frame, is ~R ~ [~ ~1 ~J As a result, the feedback force in position control mode is M F{ = k{ \u00b7 ~ R \u00b7 f7 R \u00b7 H po ' (10) where }} Po is a position vector describing the position of the origin of the object frame {o} in hand frame, k, is the coefficient of the feedback force. When the remote robot is in attitude control mode, angular velocity vector generated on the master device will control the end-effecter to adjust its attitude while its position is fixed. Given the rotation matrix ~ R which is used to describe the attitude of {o} relative to {H} ~ R = [ :: : : \u00bb, o, In Fig. 6, ~T represents coordinate transformation from {B} to {V}, IV. GENERAnON OF VIRTUAL ATTRACTIVE FORCE VT_[;R V PB] (2) B- O O O I' where {B} and {V} represent control point frame and the virtual frame respectively. Accordingly, ~T defines the coordinate transformation from object frame {o} to the frame of {B} for control points ~T = [0 goR 0 B ~ol (3) ~T defines coordinate transformation from hand frame {H} to the object frame {o} ~T = [0 f?OR 0 0 ~Hl (4) ~T is the coordinate transformation from {H} to {V} vT _ [ :; R v PI H ] \u2022 (5) H - O O O According to Fig. 6, the following equat ion can be obtained ~T . ~T = ~T f~T . (6) In Fig. 6, {V} defines a virtual frame. Just before the end effecter grasps the object, the frame {o} coincides with {II} . Meanwhile, the frame {B} , which is used to define the coordinates of control points, is consistent with the virtual frame. From (6), we have ~T = ( ~Tr . ~T . ~T, (7) where ~T and ~T are determined in the experiment. A. Feedback Force in Position Control Mode ~X _::~r ~-- j o 100 200 300 400 500 600 700 time (O.1s) the equivalent axis k and angle e which are described in { H } , will be derived, e=atan2(~(oz - as +(ax - nz)2+(nv-aS\u00bb,+ov+az-I} (12) k={kx k; kz), (13) 300 400 t ime (O" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure27-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure27-1.png", + "caption": "Fig. 27 Actuator unit of motor and ball screw", + "texts": [ + " These three joints are actuated and controlled by four cables as depicted in Fig. 26. The terminations of three cables are fixed to the foot plate at the three points around the ankle joints. The other cable goes around the knee pulley and is terminated on the upper part of the shin. By adjusting the combination of pulling force of wire cables, 3DoF can be fully controlled. As described above, the end of wire cables are fixed to the foot plate or shin. The opposite side of each cable is connected to each actuator unit consisting of a ball screw and a motor as in Fig. 27. The motor torque is converted to the linear thrust of the screw nut that generates the tension of wire cable. The actuator unit by means of ball screw has the advantage of high efficiency and no backlash. Two cables actuating the knee joint are wrapped around the pulley, the axis of which coincides with that of knee joint. By virtue of this structure, the moment arm on knee joint torque is constant regardless of knee joint angle. As a consequence, the wide range of joint movement can be achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000814_s11029-009-9095-4-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000814_s11029-009-9095-4-Figure1-1.png", + "caption": "Fig. 1. Scheme of the ini tial tu bu lar bil let.", + "texts": [ + " A clas si cal ex am ple is the man u fac ture of car tires, when a cy lin dri cal bil let is trans formed into the el e ment of a to roid al shell. In the pres - ent study, the pos si bil ity of man u fac tur ing curvilinear branches by trans form ing an un cured cy lin dri cal bil let made by sym met - ric he li cal wind ing into a curvilinear branch whose shape corresponds to the section of a circular torus is investigated. Let us con sider a cy lin dri cal shell-bil let of ra dius R and length H3 (Fig. 1), wound at an gles \u00b1b3 . For the shell, the fol - low ing re la tions are valid: ds Rd3 3 3tan= J b , dl Rd 3 3 3 = J cos b . (1) We as sume that the wind ing of the bil let is geo detic, with b3 const= . Then, as sum ing that s3 = 0 and l3 = 0 at J J3 3= 0 , we come to the equa tions for cal cu lat ing the cur rent fil a ment length l3 and the bil let length s3 l R 3 3 3 3 = -( ) cos J J 0 b , s R3 3 3 3tan= -b ( )J J 0 . At s H3 3= , the fil a ment length of the bil let is L H 3 3 3 = sin b . For a sep a rate fil a ment with an ini tial twist an gle J 03 at l L3 3= , its fi nal twist an gle is des ig nated as J L3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.10-1.png", + "caption": "FIGURE 16.10 Some common joint designs. (a) Butt joint. (b) Lap joint. (c) T-joint.", + "texts": [ + " For a given optical system (focusing lens), the location of the focal point relative to the workpiece surface determines the beam size on the workpiece surface, and this affects the bead width and depth of penetration (Fig. 16.9). Optimum location of the focus results in maximum depth-to-width ratio, that is, the maximum penetration depth and minimum bead width occur for almost the same focal position. There are various types of joint configuration that can be used in welding. However, due to the small beam size, lasers are primarily used for lap or square butt joints. The rule of thumb is that the butt gap has to be less than 10% of the thinner plate\u2019s thickness (Fig. 16.10). Otherwise there is the likelihood that most or the entire beam will pass directly through the gap, resulting in inadequate weld or no weld at all. To minimize the problem associated with the small beam size, the beam may be defocused at the joint to produce a relatively wide bead. However, the defocused beam is more sensitive to absorption by the material surface, since the power density is then low, and any small amount of power that is reflected is relatively significant. This may result in fluctuations in the energy absorbed by the workpiece" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002386_978-981-10-8597-0_24-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002386_978-981-10-8597-0_24-Figure3-1.png", + "caption": "Fig. 3 6D-PKM prototype", + "texts": [ + " The external wrench is in static equilibrium with the six-connector forces, the equation representing this is given by [ W\u0304 ] f1 \u23a1 \u23a3 s\u03021 b1 \u00d7 s\u03021 \u23a4 \u23a6 + f2 \u23a1 \u23a3 s\u03022 b2 \u00d7 s\u03022 \u23a4 \u23a6 + f3 \u23a1 \u23a3 s\u03023 b3 \u00d7 s\u03023 \u23a4 \u23a6 + f4 \u23a1 \u23a3 s\u03024 b4 \u00d7 s\u03024 \u23a4 \u23a6 + f5 \u23a1 \u23a3 s\u03025 b5 \u00d7 s\u03025 \u23a4 \u23a6 + f6 \u23a1 \u23a3 s\u03026 b6 \u00d7 s\u03026 \u23a4 \u23a6 (1) where fi(i 1,...,6) are the axial reaction forces generated in the connectors. s\u0302i(i 1,...,6) is the unit vector along the ith connector and bi is the ith connector connection point Fig. 1 6\u20136 In-parallel structure B P b0 b bb b b p 1 23 4 5 1 p0 p2p3 p4 p5 Fig. 2 6\u20136 In-parallel manipulator at the base with respect to the base frame B (see Figs. 1 and 2). Figure 3 shows the 6D-PKM designed and developed at author\u2019s laboratory. The sensors at the connectors can bypass the dead wrench load, so as to sensitize them to only active wrench. The active wrench experienced by the platform is distributed to the six legs connecting the platform and base plates. The external wrench applied is statically balanced by the six leg forces of the platform and it is important that these six leg forces are linearly independent. Though the wrench measurement fundamentally relies on Eq", + " Subsequent sections deal with those factors and experiments are conducted to show wrench guided task space trajectory control of the parallel manipulators under continuously changing manipulator configurations. In order to generate motion of the platform in response to a guiding force, it is required to measure the force sensed by the platform, and therefore the forces on the connectors or legs of the mechanism. Each leg of the mechanism is constructed by coupling a high-precision ball screw to a servomotor equipped with a position encoder (see Fig. 3). This allows the axial load to be transformed to motor torque, which corresponds to current draw, a quantity that can be measured directly. Since measurement of torque is fundamental to force control, this type ofmeasurement was first used in a software module which performs the task of bringing the mechanism to its dead-end or home position, dead-end homing. This involved jogging the motors to decrease the extension of all connectors until the moving parts on the connector and the base met, causing an abrupt increase of motor torque that would signal a stop" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003529_0954406218812119-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003529_0954406218812119-Figure5-1.png", + "caption": "Figure 5. The first-order modal: (a, c) preload of 15 kN; (b, d) preload of 35 kN.", + "texts": [ + " Fourthly, for different pre-tightening forces, both natural frequency and modal shape of the whole structure were obtained by the modal analysis. The whole structure was meshed with finite element software HYPERMESH. The modal analysis of the whole structure was obtained by ANSYS finite element analysis software. For different pre-tightening forces, natural frequency and modal shape of the whole structure were obtained by the modal analysis. Then, the experimental and simulation results were compared. The normal modal vibration shape is shown in Figure 5(a) and (b), the tangential modal vibration shape is shown in Figure 5(c) and (d), wherein the pre-tightening forces of 15 kN and 35 kN are taken as examples. The natural frequencies of the first order of the whole structure under different preloads were obtained by the modal analysis. The results are shown in Table 1. The comparison between experimental identification and simulation results is shown in Figure 6. Comparing natural frequencies for different pre-tightening forces, it can be seen that the maximal relative error does not exceed 5%. The results also show that the model established can correct the normal and tangential stiffness and damping of joint surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000996_978-1-4020-6114-1_15-Figure15.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000996_978-1-4020-6114-1_15-Figure15.1-1.png", + "caption": "Fig. 15.1. Six unmanned vehicles perform a formation tracking task.", + "texts": [ + " law, and in [2] [12] [16] through the use of dynamic feedback linearization. The backstepping technique for trajectory tracking of nonholonomic systems in chained form was developed in [6] [10]. In the special case when the vehicle model has a cascaded structure, the higher dimensional problem can be decomposed into several lower dimensional problems that are easier to solve [17]. An extension to the traditional trajectory tracking problem is that of coordinated tracking or formation tracking as shown in Figure 15.1. The problem is often formulated as to find a coordinated control scheme for multiple unmanned vehicles that forces them to maintain some given, possibly time-varying, formation while executing a given task as a group. The possible tasks could range from exploration of unknown environments where an increase in numbers could potentially reduce the exploration time, navigation in hostile environments where multiple vehicles make the system redundant and thus robust, to coordinated path following. Detailed information may be found in recent survey papers [1] [21]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003334_icsengt.2018.8606399-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003334_icsengt.2018.8606399-Figure2-1.png", + "caption": "Fig. 2. The structure and frame of hexacopter", + "texts": [ + " SYSTEM MODELING The Hexacopter UAV motion consists of three force and three moment component which form six Degree of Freedom (6 DoF). As the dynamics systems, six rotors speeds are the systems input and position (x, y, z) and attitude (\u03a6, \u03b8, \u03a8) are system outputs. The combination of rotor\u2019s speed variation generate some motions, such as: throttle, roll (rotation arround x axis), pitch (rotation arround y axis) and yaw (rotation arround z axis), which are described in equation (1) as U1, U2, U3 and U4 respectively. Figure 2 ilustrated the schematic structure of hexacopter and the rotational directions. There are two reference that necessary to describe the hexacopter motion, which are earth inertial frame and body fix frame. Equation (2) denotes the transformation of the body inertial frame to earth inertial frame [9,10]. = (1) = \u2212 + ++ \u2212 +\u2212 (2) = 10 \u22120 (3) The dynamic model of hexacopter presented in equation (5) based on the equation of motion some forces and torques applied in it, that give the translational and three rotational motion for hexacopter with respect to the body frame, which is described in equation (4)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000120_indin.2008.4618144-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000120_indin.2008.4618144-Figure1-1.png", + "caption": "Fig. 1 The prototype of the system", + "texts": [ + " A novel approach that includes three processes is proposed in the paper. Firstly, the transformation between the robot wrist and the contact position sensor tool is calibrated. Secondly, the transformation between the robot tool and the external axle worktable is calibrated by the robot and the contact position sensor joint measurement system. Finally, Particle Swarm Optimization (PSO) Algorithm is adopted to optimize the key parameters to improve the accuracy of the total system. The robot scanning system with an external axle is shown in Fig.1. The system consists of an external axle and a six-degree freedom robot with an inductance contact position sensor fixed on its wrist flange. The robot coordinates is shown in left part of Fig.2. The frame {B}, located on the center of the robot\u2019s base, is the reference coordinates of the system. The center of the flange at the bottom of the robot is defined as frame {W}. The zero reading point of the sensor is defined as the 1 The work was supported by national natural science foundation of China (50575029) and (60674061) 2 Corresponding author: e-mail: zima@newmail" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002287_s00170-018-2422-y-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002287_s00170-018-2422-y-Figure3-1.png", + "caption": "Fig. 3 Bent workpiece enclosing design model. (a) \u03b1f >\u03b1, (b) \u03b1f <", + "texts": [ + " Subsequent to the prediction of springback using Gardiner\u2019s estimation formula coupled with experiment adjustments, \u03b1f (final bend angle) can be obtained and will be used to calculate the optimal thickness to. The optimal thickness to is the minimum thickness of workpiece that can enclose the design model; therefore, it can be used as the baseline for stock material selection. Depending on the relation of final bend angle \u03b1f and design bend angle \u03b1, there are two cases that have impact to the calculation of optimal thickness to: (1) \u03b1f is greater than \u03b1 and (2) \u03b1f is less than \u03b1. Figure 3 illustrates the difference of these two cases schematically. In the author\u2019s previous work [24], the calculation of optimal thickness to was developed when \u03b1f is greater than \u03b1, and was used to guide the selection of workpiece material for angular accuracy enhancement. In this subsection, the method in the author\u2019s previous work [24] for calculating optimal thickness to when \u03b1f is greater than \u03b1 is first recapitulated, and then it is extended to the case where \u03b1f is less than \u03b1. (1) \u03b1f is greater than \u03b1 In the author\u2019s previous work [24], the design model is fitted into the post-bending workpiece model, and the optimal thickness is analytically found to completely enclose the design model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001545_978-3-319-05633-3_8-Figure8.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001545_978-3-319-05633-3_8-Figure8.1-1.png", + "caption": "Fig. 8.1 Left: schematic representation of the spring mass model of running. The centre of mass trajectory can be roughly described by that of the white circle at the great trochanter. Right: simplified geometry of the spring-mass system during the support phase. L: lower limb length at initial contact; h: angle swept by the system during the forward motion for each half of the support phase. Assuming a constant running velocity (v in m s\u22121), the horizontal displacement of the centre of mass during the support phase is equal to v tc/2, with tc the contact time in s", + "texts": [ + " We will then show examples of use and applications, and discuss the main 1Although running involves motions of the body in all three planes of space, we will limit our approach in this chapter to the sagittal plane of motion. 2See for instance the running robot Phides from the Delft University of Technology: http://www. 3me.tudelft.nl/en/about-the-faculty/departments/biomechanical-engineering/research/dbl-delftbiorobotics-lab/bipedal-robots/. 3See for instance the M.I.T Cheetah Robot: http://biomimetics.mit.edu. technologies that can be used to measure the mechanical inputs of this simple method. The SMM (Fig. 8.1) models the runner as a mass (equal to the total body mass) loading a linear spring that represents the lower limb (thigh, leg and foot segments, hip, knee and ankle joints). At each step, as the runner lands on the ground, the mass compressed the spring during its downward motion, and then the spring extends in the second part of the ground contact phase, before the subsequent aerial phase. Contrary to rebound vertical jump and hopping (see Chap. 6), the SMM for running implies an oscillation forward during this compression-extension cycle", + "3): 4Although it has an indirect relationship, this stiffness is distinct from what is sometimes referred to as \u201cmuscle stiffness\u201d or \u201cjoint stiffness\u201d. The latter usually indicate a loss of range of motion and the possibly associated pain. F\u00f0t\u00de \u00bc Fmax sin p tc t \u00f08:1\u00de Knowing the vertical GRF data and the runner\u2019s body mass m allows one to compute the vertical displacement of the center of mass during the support phase using the laws of motion (Cavagna 1975). Then, using the simplified stance phase geometry shown in Fig. 8.1, the change in leg length during contact can be computed (McMahon and Cheng 1990; Farley and Gonz\u00e1lez 1996). Then, as shown in Fig. 8.3, a plot of the vertical GRF against the leg length changes shows a pretty linear trace. On this basis, the SMM for running has been used to analyze running locomotion mechanics in humans, but also in biped and quadruped animals (Alexander 1991, 2003, 2004; Farley et al. 1993; Kram and Dawson 1998). Definitions and Assumptions Since the lower limb compression and length change follow an overall out-of-phase behavior during the stance phase, a simple way to compute the lower limb stiffness (kleg) is to divide the maximal compression force by the corresponding maximal leg length change (Farley and Gonz\u00e1lez 1996)", + " \u2022 the SMM and the computation of stiffness that will be detailed in the next paragraph are based o the fact that the maximal compression force (Fmax) and the minimal length of the lower limb (that corresponds to ymin in Fig. 8.3, and that is used to compute the leg length change \u0394L) are reached simultaneously. As shown in the typical example of Fig. 8.3, the delay is very short between these two instants (Silder et al. 2015). \u2022 the two main assumptions that are made when using the SMM for running are related to the geometrical representation of the lower limb motion during the stance (Fig. 8.1). First, the point of force application onto the ground is assumed to be fixed during the stance. This is not the case in reality, since the center of pressure moves by about 10\u201320 cm under the foot during contact. This has been shown to influence the stiffness computations (Bullimore and Burn 2006) and a correction could be applied, with an estimation of the length of this \u201cforward translation of the point of force application\u201d based on running velocity (Lee and Farley 1998). The important point here is that, in case of intra-subject comparisons or inter-subjects comparisons using the same computation method (training process or other similar investigations), this assumption does not challenge the results obtained. Second, the angular sector swept by the lower limb during running is assumed to be symmetrical around the vertical mid-stance position (Fig. 8.1), and the length of the lower limb at foot strike L is assumed to be equal to the reference value measured from the great trochanter to the ground in anatomical position. This simplification has been discussed and video analyses confirmed that the lower limb length at foot strike is overestimated in the classical SMM (Arampatzis et al. 1999). The main point in our opinion is that this assumption of a symmetrical angle swept during the support phase makes the SMM invalid in accelerated/decelerated running, and during uphill/downhill running", + " Although synchronizing a high-frequency video analysis with GRF data can lead to accurate measurement of center of mass position over time, both Fmax and \u0394y can be derived from the sole GRF signal using a method detailed by Cavagna (1975). Briefly, this method uses the laws of mechanics to compute vertical acceleration of the center of mass over time, and then velocity and position are obtained by integrating this acceleration signal (Fig. 8.4). Then, \u0394L is computed from the lower limb geometry presented in Fig. 8.1 as follows (Farley et al. 1993; Farley and Gonz\u00e1lez 1996; Farley and Ferris 1998): DL \u00bc L ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L2 vtc 2 2 r \u00feDy \u00f08:4\u00de Note that L can be estimated from subject height H using the anthropometrical model of Winter (1979): L \u00bc 0:53H \u00f08:5\u00de Ground reaction force data are typically measured with force-plate systems embedded into the running ground (e.g. Slawinski et al. 2008b; Rabita et al. 2011) or with instrumented treadmills (e.g. Dutto and Smith 2002)", + " Assumptions of the spring-mass model. The simple method presented here is based on the use of the spring-mass model of running. Therefore, all the limitations and assumptions associated with the use of this model (see Sect. 8.2.4) also apply to the use of this method. Running conditions: acceleration, speed and slope. The linear spring-mass model is based on a geometrical consideration of the lower limb during the stance phase that is mainly characterized by a symmetry of the angle swept before and after mid-stance (Fig. 8.1) (Blickhan 1989; McMahon and Cheng 1990; Farley and Gonz\u00e1lez 1996; Bullimore and Burn 2006). Consequently, all experimental conditions that clearly induce a running pattern in which this basic assumption of the model is infringed lead to questionable data. Although all inputs of the simple method used may be measured in any running condition, the spring-mass variables computed (e.g. vertical and leg stiffness) do not make much sense if the basic postulates of the model are not respected. These conditions include in particular non-constant velocity (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000779_icit.2009.4939600-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000779_icit.2009.4939600-Figure2-1.png", + "caption": "Fig. 2. The full 3D model of the stator end windings.", + "texts": [ + " So the edge element method is extensively used in motor magnetic field analysis. As edge element method can not directly calculate the inductance of solid source conductor, it can not realize direct circuit field coupling. In this paper, a novel methodology for coupling edge element analysis with circuit simulation is presented. The distribution of eddy current field in the iron core and core loss is calculated. For a flameproof induction motor, the typical structures of the end region are shown in Fig. 1. The full 3-D model of the stator end windings are shown in Fig. 2. Fig. 3 shows the circuit diagram of the flameproof motor. To impose loads on the 3-D finite element model expediently, the circuit-field coupled approach is adopted. Two basic approaches to coupling finite element analysis with circuit simulation exist. One is direct coupling approach [5] where the field and circuit equations are coupled directly together and solved simultaneously. The other approach is indirect coupling where the finite element analysis and circuit simulator are treated as separate systems in a step-by-step process with respect to time, while they exchange coupling variables in each step [6]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003115_ecce.2018.8557727-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003115_ecce.2018.8557727-Figure2-1.png", + "caption": "Fig. 2. 3D thermal model.", + "texts": [ + " The mutual coupling between the two healthy 3-phase sets and one faulty 3-phase set will result in an magneto-motive force (MMF) offset component in the region occupied by 3-phase set ABC [14]. Therefore, the location of the SC turn will affect the flux linkage and consequently the circulating current and resultant copper loss. It has been investigated in [15] that when the SC turn located at slots B2 and B4 which are marked by the two black quadrangles shown in Fig. 1 (a), the SC current and copper loss are the highest. The subsequent analysis is focused on this worst case. III. 3D THERMAL MODEL The 3D thermal model is shown in Fig. 2 where different components are indicated. The 3D model encompasses 12 slots and one half of the machine axial length. To reduce modelling complexity, the end winding is simplified in the 3D thermal model as straight winding segment with the same equivalent length as those in the prototype machine. The 3D FE model together with the schematic heat transfer network with the ambient and cooling circuit shown in Fig. 3 forms the complete thermal model of this motor in JMAG [16]. Additionally, the end winding is covered by the potting in the 3D thermal model which is not drawn in Fig. 2. As observed from Fig. 3, the radiation heat transfer is negligible in the study because of the small temperature difference between the machine surfaces and ambient. Besides, it is important to compute thermal properties and parameters, such as conductivity, contact thermal resistance, convection coefficients and thermal capacitance, to build the 3D thermal model. The commercial software package, Motor-CAD, [17] as well as empirical equations [1] are used to compute all these parameters. The thermal conduction resistance Rcond is given by [1]: cond LR kA = (1) where L is the length in the thermal conduction path; A is the surface area and k is the thermal conductivity of the material", + " The prototype PMASynRM is tested under two operations, healthy condition and one turn short circuit with 3-phase terminal short circuit. In both cases, the machine operates at the base speed of 4000 rpm with load current in healthy phases being set to 120A for maximum torque per Ampere (MTPA) operation. The thermal tests are performed for 2 hours for each case. The machine temperatures under these two conditions are predicted by the 3D thermal model and the results are compared with the test results. As the temperature distribution is the same in each 3-phase set under healthy condition, the 3D thermal model in Fig. 2 is adopted in simulation. The resultant temperature distribution shown in Fig. 8 is almost symmetric in each 3-phase set, and the overall temperature in the end winding is higher than that in the active winding. However, the temperatures of active winding in each slot are slightly different due to the fact that the copper loss assigned to the end-winding segment associated with each slot are different, being proportional to the real end winding layout. Significant temperature differences in each slot are seen in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000687_ichr.2009.5379565-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000687_ichr.2009.5379565-Figure2-1.png", + "caption": "Fig. 2. Mass and center of gravity used for the dynamic modeling", + "texts": [ + " On figure 1, the right leg (coordinates ql and q2) is supposed to touch the ground and the left leg (coordinates q4 and q5) is in swinging phase and behind the hip. The dynamic model of the robot is given by the following equation : D(q)q+H(q,q)-Q(q) = Br+Ac1 (q)Fr+Ac2(q)FI (1) where q = [ql q2 q3 q4 q5 x zf is the absolute coordinate vector, q = [tiI q2 q3 q4 q5 X i ]T is the speed vector, D(q) is the mass matrix, H(q,q) is the centrifugal and Coriolis forces vector, r the torque vector on the joint (positive direction is indicated on Fig.2), B torque connection matrix, Q(q) is the gravity vector, Ac1 (q)F; and Ac2 (q).Fl are the generalized constraint forces which act on right and left foot respectively. The expression of the torque vector is due to the four actuators (on the h* and on the knee) and is given by r = [r1 r 2 r 3 r 4] . The expression of the torque connect ion matrix is: -1 0 0 0 1 -1 0 0 0 1 1 0 B= 0 0 -1 1 (2) 0 0 0 -1 0 0 0 0 0 0 0 0 The inertia matrix of the robot is a symmetrical, definite and positive matrix and is given by: Dn D12 0 0 0 D16 D17 D12 D22 0 0 0 D26 D27 0 0 D33 0 0 D36 D37 D(q) = 0 0 0 D44 D45 D46 D47 0 0 0 D54 D55 D56 D57 D16 D26 D36 D46 D56 D66 0 D 17 D27 D37 D47 D57 0 D77 (3) The explicit expressions of the coefficients of these various vectors and matrices are given in appendix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000087_12.819551-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000087_12.819551-Figure1-1.png", + "caption": "Fig 1 gear head-cutter", + "texts": [ + " The family of a generating surface (the cone surface) is generated while the cradle and gear, which is being generated, perform related rotations about the fixed axes and the axes of the being generated gear. *caoxuemei-2002@163.com Fourth International Symposium on Precision Mechanical Measurements, edited by Yetai Fei, Kuang-Chao Fan, Rongsheng Lu, Proc. of SPIE Vol. 7130, 71300F \u00b7 \u00a9 2008 SPIE \u00b7 CCC code: 0277-786X/08/$18 \u00b7 doi: 10.1117/12.819551 Proc. of SPIE Vol. 7130 71300F-1 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 06/26/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx Fig1 shows the cone of gear head-cutter. The derivation of the gear tooth surface is based on the following procedure: Step1: the cone surface is represented as follow: g g 2 g e g g g g 2 g g 2 ( sin )cos ( , ) ( cos )sin cos r u r u r u u \u03b1 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u23a1 \u23a4\u2212 \u23a2 \u23a5= \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 ur (1) Here: gu and g\u03b8 are the surface coordinates; 2\u03b1 is the blade angle; gr is the radius of gear head-cutter. Step2: by the equation of meshing and the coordinate transformation, the gear tooth surface is represented as follow: 2 2 g g( , )r r \u03b8 \u03c6= r r (2) Here: g\u03c6 is the rotation angle of cradle in the process of gear generation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure6-1.png", + "caption": "FIGURE 6 Gear tooth radial shift", + "texts": [ + " 1\u00a0:\u00a0Design profile 2\u00a0:\u00a0Actual profile 3\u00a0:\u00a0Mean profile 1a\u00a0:\u00a0Design profile trace 2a\u00a0:\u00a0Actual profile trace 3a\u00a0:\u00a0Mean profile trace 4\u00a0:\u00a0Origin of involute 5\u00a0:\u00a0Tip\u00a0point 5\u00a0\u2212\u00a06\u00a0:\u00a0Usable profile 5\u00a0\u2212\u00a07\u00a0:\u00a0Active profile C\u00a0\u2212\u00a0Q\u00a0:\u00a0Base tangent length to point\u00a0C \u03beC\u00a0:\u00a0Involute roll angle to point\u00a0C Q\u00a0:\u00a0Start of roll angle A\u00a0:\u00a0Tooth\u00a0tip\u00a0or start of chamfer C\u00a0:\u00a0Reference point E\u00a0:\u00a0Start of active profile F\u00a0:\u00a0Start of usable profile LAF\u00a0:\u00a0Usable length LAE\u00a0:\u00a0Active length L\u03b1\u00a0:\u00a0Evaluation range LE\u00a0:\u00a0Base tangent lenth to start of active profile F\u03b1\u00a0:\u00a0Total profile deviation ff\u03b1\u00a0:\u00a0Profile form deviation fh\u03b1\u00a0:\u00a0Profile slope deviation Figure 6 shows an involute tooth profile with a vertical shift after assembly overlaid on an involute tooth profile of an unassembled gear for comparison. The distance between the unassembled and the assembled profile as measured along the normal to the profile increases with increase in the roll angle and is a function of the radial deformation and the pressure angle at the point of measurement. The tooth profile curvature and the associated pressure angle change gradient with gear diameter, is greater near the root and decreases towards the tip with increase in the radius of curvature and the associated roll angle", + " As expected, the profile deviation due to radial distortion increases from the fillet to the tip. Figure 7 shows the expected profile measurements \u00a9 2018 SAE International. All Rights Reserved. before and after assembly based on the calculation procedure developed. Figures 8 and 9 are the actual measured profiles of the gear before and after assembly and the change in profile slope is obvious. The change of profile slope deviation between the measured tooth profiles before and after assembly is accurately calculated by the calculation procedure discussed earlier. Figure 6 shows the change in tip diameter due to radial deformation. The increase in the tip diameter is approximately equal to the change in the inner diameter of the gear due to interference as the deformation in circumferential and axial directions is minimal. The change in tip diameter is symmetric and same around the gear. Maximum interference assembly has greater increase in tip diameter than minimum interference assembly, which in turn results in greater increase in tip diameter than minimum interference assembly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001674_978-3-319-76138-1_3-Figure3.14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001674_978-3-319-76138-1_3-Figure3.14-1.png", + "caption": "Fig. 3.14 For the zonotope of the available wrench set Q, we can check if the ellipsoid E with the desired wrenches is fully contained", + "texts": [ + " ,wn} of wrenches can be generated, every wrench wP in the convex hull of the wrenches wi can also be generated. \u2022 Thus, to check if a hypercube of wrenches can be generated, it is sufficient to check all of its corners. Still, in n-dimensional space, the hypercube has 2n corners but checking all corners is straightforward. Therefore, for the most general 3R3T robot, one has to consider 64 corners. \u2022 In a similar way, one can check if every wrench wP within an ellipsoid E of wrenches can be generated by the robot. To do so, one has to check only 2m characteristic points on the surface of the ellipsoid (Fig. 3.14). The characteristic points are determined on the surface of the ellipsoid E through the normals of the hyperplanes which generate the zonotope [62]. \u2022 The zonotope defining the wrench set Q can be constructed from the Minkowski sum of m generating lines (Fig. 3.15). \u2022 Once the zonotope forming the wrench set has been calculated, it is simple to check for wrench-feasibility for any given wrench by testing if the wrench wP or a set of wrenches Q is fully enclosed by the zonotope. \u2022 If the zonotope degenerates, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000007_amr.44-46.829-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000007_amr.44-46.829-Figure8-1.png", + "caption": "Fig. 8 K6 type bogie and the forces transferred from the side frame acting on the adapter", + "texts": [ + " The closer to axle side, the larger the equivalent stress amplitude. 2. Maximum equivalent stress amplitude appears at the outer ring inboard side close to the seal seat. K6 Type Bogie. At the same time, 353130B bearings are also used on the same axle weight freight cars with K6 type bogie. Number of bearings on K6 type of bogie is actually more than those on K5 type bogie. But up to now, no failure has occurred on this bogie. An EPFE analysis is carried out on the bearing on K6 type of bogie to understand the difference of two kinds of bogie. As shown in Figure 8, this kind of bogie has a structure of down crossing bars to control the vehicle waving laterally right or left. Force from the side frame of bogie is transferred to bearing through rubber acting directly on the adapter. By the wheel-track contact force spectrum under same piece of railway line and same speed, results of EPFE analysis are displayed in Table 3 with respect to the selected inspection positions shown in Figure 7. It verifies that 1. The equivalent stress amplitudes show heterogeneously axial distributions for rollers and inner- and outer-rings of the bearing for K6 type of bogie" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001211_6.2009-5885-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001211_6.2009-5885-Figure1-1.png", + "caption": "Figure 1. Desired heading in the presence of wind.", + "texts": [ + " Some control systems on the other hand use current relative position error and current relative velocity error to obtain control commands for the follower aircraft12. Again this needs communication to be instantaneous. Apart from that, they set their desired heading as the direction of the desired velocity vector, , which is obtained from the relative position error, , and the velocity of the leader, . But maintaining heading along will not be the right course of action in the presence of winds as can be inferred from Fig. 1. The heading will have to be maintained at an angle, , to the direction of , and this angle would depend on the speed and direction of the wind (which is continuously varying) and of the aircraft itself. American Institute of Aeronautics and Astronautics 3 To overcome these problems, our approach requires that the leader communicate 3 pieces of information to the follower aircraft: 1. Current leader position, in GPS coordinates: , 2. Estimated leader position after 2 seconds, in GPS coordinates: , 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001617_s40799-018-0232-7-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001617_s40799-018-0232-7-Figure2-1.png", + "caption": "Fig. 2 3D model of a testing bench section in AutoCAD\u00ae 2016 student version", + "texts": [ + " Consequently, this model will provide the basic geometry for the finite element software. The first stage is modelling the testbed in AutoCAD 2016\u00ae student version. For that purpose, it is necessary to review the main topics of engineering drawing such as: 1) Fundamentals of AutoCAD, 2) Drawing views, 3) Auxiliary views, 4) Crosssectional views, 5) Dimensioning, 6) 3D Modelling, 7) Projections, and finally 8) Auxiliary views from 3D models. The resulting models need to be saved in a neutral format such as IGES or STEP to be ultimately compatible with ANSYS\u00ae software. Figure 2 shows a section of the testbed drawn in AutoCAD. Once the testbed model is finished, the finite element analysis is carried out. In order to carry out this stage, the students must undertake a 20-h basic training in ANSYS Workbench involving the following topics: 1) Fundamentals of ANSYS APDL and ANSYS Workbench, 2) Plane stress, 3) Solid modelling, 4) Contact and large displacements and 5) Fundamentals of geometry importation and rigid dynamics. Once the training is finished, the students have covered the topics related to drawing geometries, assignment of materials, discretisation of models, symmetry, sub-modelling, types of contact, non-linearities, parameterization and data processing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003093_978-981-13-1903-7_27-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003093_978-981-13-1903-7_27-Figure1-1.png", + "caption": "Fig. 1 Oversimplified diagram of the bifurcated artery with stenosis", + "texts": [ + " The influence of suction and injection on saturated micropolar fluid flow through the porous medium has been investigated by Ram Reddy and Pradepa [7]. This article deals with the flow of couple stress fluid through a bifurcated artery by treating walls as porous plates. Consider the flow of blood in a bifurcated artery withmild stenosis in its parent artery by treating the walls of the artery as porous plates. Blood is treated to be couple stress fluid. The stenosis and bifurcation of the artery are taken to be in an axisymmetric manner as shown in Fig. 1. The cylindrical polar coordinate system is considered for frame of reference in which z-axis is taken along the central line of the parent artery. In order to eliminate the flow separation zones, deflection is initiated at the start of the lateral junction and the apex. The governing equations for the flow of incompressible couple stress fluid under the influence of uniform transverse magnetic field in the porous medium is given by \u2207 \u00b7 q 0 (1) \u03c1(q \u00b7 \u2207)q \u2212\u2207 p + \u03bc\u22072q \u2212 \u03b7\u22074q \u2212 \u03bc k q + J\u0304 \u00d7 B\u0304 (2) where \u03b7\u2014couple stress viscosity parameter, \u03bc\u2014dynamic viscosity of blood, \u03c1\u2014density of blood, q\u2014velocity vector, B\u0304\u2014strength of magnetic field, J\u0304\u2014current density, \u03ba\u2014permeability parameter of porous medium and the body force and body moments are neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001566_msf.916.85-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001566_msf.916.85-Figure1-1.png", + "caption": "Fig. 1. SLM process optimization. Cross-section melt pool for separate laser passages in Ti+15%vol TiC system: \u0430) substrate heating to 200 0\u0421; b) room temperature; c) top view, left - heating to 200 0\u0421, right \u2013 to room temperature; d) appearance of the 3D gradient parts after the SLM. Cube sizes 5\u04455 mm.", + "texts": [ + ", RF) with - 50+20 \u00b5m particle size; (b) nano sized titanium carbide powder graded 99+wt% TiC, (40-60 nm, cubic) from (US Research Nanomaterials Inc., Houston, USA). The powders were premixed in different proportions: 85 and 15; 90 and 10; 95 vol. % of Ti and 5vol. % of titanium carbide, respectively. Our experimental setup was described earlier in [11]. The hatching distance was equal to the laser beam diameter db - 70 mm, layer thickness H was ~ 0.2 mm. The layers were made out of premixed Ti + nano TiC powders on a substrate by the following scheme (Fig. 1): the first ten layers were of titanium with 5 %vol. TiC, the second ten consisted of 90% Ti + 10%vol. TiC, and the third layers - of 85% Ti + 15%vol. TiC. The laser scan velocity V ranged from 0.5 to 10 cm/c, laser power P - from 10 to 100 W. Two regimes of manufacturing have been studied \u2013 with the substrate additional heating up to 200 0C and without it. Each second layer was formed on the bottom layer after its turning by 90 degrees (L \u2013 longitudinal, T - transversal). The laser melting process was conducted in Ar gas-filled chamber in order to protect the samples against oxidation and nitration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.21-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.21-1.png", + "caption": "FIGURE 16.21 Schematic of the multiple-beam laser welding concept. (From KannateyAsibu Jr., E., 1991, ASME Journal of Engineering Materials and Technology, Vol. 113, pp. 215\u2013221.)", + "texts": [ + " This slows down the production rate, since it introduces an additional step into the process. Second, since the entire workpiece is often heated, energy is wasted, and also the heat applied may affect other portions of the workpiece. Finally, conventional preheating and postweld heat treatment are often either too expensive or impractical to use since the part may be complex in shape or bulky. The drawbacks of traditional preheating and postweld heat treatment can be eliminated by using the multiple-beam configuration, which, in its simplest form, is illustrated in Fig. 16.21. In this configuration, there are two beams. One of the beams (the minor beam) is defocused, leads, and preheats the joint, while the major beam follows and welds the material. In another configuration, the defocused beam would follow the major beam, resulting in postweld heat treatment. Either arrangement reduces the cooling rates, thereby reducing or preventing the formation of hard microstructures. This is especially useful in welding high hardenability materials such as high carbon and/or alloy steels and some titanium alloys" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002940_icsedps.2018.8536040-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002940_icsedps.2018.8536040-Figure9-1.png", + "caption": "Fig. 9 Voltage vector v2 to debilitate the stator flux vector.", + "texts": [ + " \u03b8 \u03bb\u0394+\u03bb * \u03bb\u0394\u2212\u03bb * \u03b8\u2212\u03a0 4 Assuming *k \u03bb \u03bb\u0394= and * x'x \u03bb = )cos()k(x)k(x)k( ** \u03b8\u03c0 \u03bb\u03bb +\u2212\u2212\u2212+=\u2212 4 31211 2 2 2 2 Substituting and solving 04 4 12 =\u2212+\u2212+ k)sin()k('x'x \u03b8\u03c0 By comparing it with quadratic equation 02 =++ cbxax and apply it in the formula a acbb 2 42 \u2212+\u2212 We obtain )sin()k()(sin)k(k'x \u03b8\u03c0\u03b8\u03c0 ++\u2212+++= 4 1 4 14 22 (18) By applying maclaurin series in the above equation, we obtain 1 4 4 22 2 2 2 + + +\u2212 += )(sin )(sin(k k'x \u03b8\u03c0 \u03b8\u03c0 The above equation can be compared to the below equation 12 ++= akk'x Since the k is of small value, the flux developing vector of normalize length is obtained as In the second case, we have presumed that the stator flux modulus has come to the upper limit of allowed pemitting value of \u03bb\u0394\u03bb + * in the sector 1. Therefore from Fig. 9, v2 is choosed to decrease the stator flux modulus. Similar calculations in OA\u2019B\u2019 triangle, which gives the normalized distance of flux-debilitating vector as )cos( k'y \u03b8\u03c0 + = 4 2 (20) Therefore by putting the values obtained from equation (19) and (20) we can derive the probabilities as )cos()sin( )cos( 'y'x 'x)(Pv \u03b8\u03c0\u03b8\u03c0 \u03b8\u03c0 \u03b8 +++ + = + = 44 4 1 (21) )cos()sin( )sin( 'y'x 'y)(Pv \u03b8\u03c0\u03b8\u03c0 \u03b8\u03c0 \u03b8 +++ + = + = 44 4 1 (22) From (19), (20), (21), and (22) .Table II, the linear speed can be derived as \u03b8 \u03b8\u03b8\u03b8 cos Vv)(P)(v)(PV dc pvpvp 2 21 21 =+= (23) Therefore, by integrating the stator flux vector, we can obtain the rotational speed as * dc * dc avg V.d )cos( V \u03bb \u03b8 \u03bb\u03b8\u03c0 \u03c9 \u03c0 \u03c0 560 2 2 4 4 \u2245 = \u2212 (24) The equation no. 24 shows the equation of average and its relation with stator flux modulus. By directly putting the value of speed and DC terminal voltage, value of stator flux modulus can be obtained. This modulus can be used as a reference value [1,12]. V. SIMULATION PARAMETER VI. SIMULATION RESULTS Fig.9 Stator current vs. time Fig. 10 Stator flux modulus vs. time VII. ANALYSIS OF SIMULATION RESULTS Fig. 9 shows stator current. The two components of currents i.e. d-axis and q-axis component is shown. There is 90 degrees phase shift between two currents. Fig. 10 shows the modulus of stator flux. The band of stator flux modulus is very small. Fig. 11 shows sectors for generating pulses. The counterclock wise rotation is considered. In the fig. four sectors are generated which are shown in different levels. Fig. 12a shows waveform of torque. As per equation 15 and Sr. no Parameters Values 1 Stator resistance 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.4-1.png", + "caption": "Fig. 11.4 Brush model", + "texts": [ + " These forces are applied to a point that is located at a length t along the X-axis measured from point O under the wheel center. ttrail is called the pneumatic trail. When F is the total force of a tire, the relations among Fdrag x , Fx, FCF y and Fy are F2 \u00bc F2 x \u00feF2 y \u00bc Fdrag x 2 \u00fe FCF y 2 Fdrag x \u00bc Fy sin a\u00feFx cos a FCF y \u00bc Fy cos a Fx sin a: \u00f011:2\u00de When the force is measured by a force transducer on a drum tester, Fdrag x and FCF y are measured. When the force is measured by a force transducer on the wheel, Fx and Fy are measured. (1) Fundamental equations for the solid tire model with a brush model Figure 11.4 shows the solid tire model with a brush model consisting of bristles called tread elements, the properties of which are determined by the shear spring rate of tread discussed in Sect. 7.1. When bristles roll in the contact area, they are normal to the contact surface at the leading edge. When the force at the tread element is less than the maximum frictional force, the tread element adheres to the contact surface. If the force is more than the maximum frictional force, the tread element slips on the contact surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003312_imece2018-86959-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003312_imece2018-86959-Figure1-1.png", + "caption": "Fig. 1 schematic diagram of creep feed grinding", + "texts": [ + " The elemental composition of Ni3Al intermetallics IC10 is described in Table 1. The physical properties of Ni3Al intermetallics is presented in Table 2. The size of the sample was 10 mm\u00d710 mm\u00d720 mm. on CHEVALIER FSG-B818CNC creep feed grinder. The grinding method chosen for these experiments was down grinding. In order to make the grinding wheel surface topography consistent during each experiment, the wheel is dressed after each grinding. The schematic diagram of grinding process is shown in Fig. 1. The grinding wheel used for IC10 grinding is a white alumina (WA) and pink fused alumina (PA) mixed grinding wheel with a grain size of 80 mesh. The typical grinding parameters of creep feed grinding are grinding depth 0.02~1mm, workpiece speed 30~300mm/min and wheel speed 10~35m/s. The experimental parameters designed according to the typical range of grinding parameters are shown in table 3. 2 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002373_s0031918x18080057-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002373_s0031918x18080057-Figure5-1.png", + "caption": "Fig. 5. Dependence of the component averaged over the computational region on H for the case of two defects, L = 100 nm. Corresponding distributions of magnetization in the plane z = Lz are analogous to those shown in Fig. 4b.", + "texts": [ + " 1 at which the transitions from the loop of one type to that of another type occur depend on the size and, probably, parameters of the magnetic inhomogeneity (this question is a subject for further study). Then, a series of numerical experiments was carried out in which we studied the dependence of the magnetization curves on the characteristic size of the inhomogeneities (side a of the square in the xy plane, see Fig. 1); the inhomogeneities are located at the distance L = 100 nm. Results are shown in Figs. 9\u201312. The comparison of the hysteresis loops in Figs. 9 and 10a with Fig. 5 demonstrates that the width of the loops decreases with decreasing size. At the same time, the position of the vortex core with respect to the center of the defect hardly changes (compare Figs. 9b, 10b, and 4b). If the size of the defect decreases to a cer- 19 No. 8 2018 zm zm tain critical value less than the characteristic size of the intrawall vortex (Fig. 11b), then the transition to the two-loop configuration of the magnetization curve occurs and, in this case, the width of the right-hand hysteresis loop becomes almost zero (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003718_9781119313649.ch8-Figure8.23-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003718_9781119313649.ch8-Figure8.23-1.png", + "caption": "Figure 8.23 (A) ECL of the immunosensor at various concentrations of CA199 (U ml\u22121): (a) 0.002, (b) 0.01, (c) 0.05, (d) 0.5, (e) 5, (f ) 50. (B) Relationship between ECL and the CA199 concentration, each point being the average of 10 measurements. (C) The selectivity of the ECL immunosensor. (D) ECL\u2013potential curves of (a) pure C-dots, (b) pPtPd@C-dots and cyclic voltammogram (inset) of pPtPd@C-dots on the electrode in pH 7.4 phosphate buffer saline (PBS) containing tissue plasminogen activator (TPA) [154]. Source: Reprinted with permission from Elsevier, Copyright 2014.", + "texts": [ + " In addition to detecting charge in SETs, C-dots have also been applied to the construction of electronic sensors for the detection of humidity and pressure [152]. ECL emission was also observed from C-dots [153\u2013157]. Yang et al. designed a sandwich-type Carbon Nanodot Composites: Fabrication, Properties, and Environmental and Energy Applications 257 ECL immunosensor by using gold\u2013silver nanocomposite-functionalized graphene (GN-Ag\u2013Au) as the sensor platform and C-dot-functionalized porous Pt/Pd nanochains (pPtPd) as the signal amplifier for the detection of a tumor marker (see Figure 8.23) [154]. Lu et al. constructed an ECL sensor via an electrochemiluminescence resonance energy transfer (ERET) between C-dots and gold nanoparticles (Au NPs) for DNA damage detection [155]. By combining intense ECL of C-dots and an aptamer technique, Lu et al. proposed a novel ECL aptamer sensor for measuring ATP [156]. In addition, Dong et al. presented an ECL sensor for the detection of H2O2 using hydrazide-modified C-dots [157]. On the other hand, the electrocatalytic performances of C-dots have been gradually recognized, allowing the use of C-dots as nanoprobes for electrochemical biosensors [158\u2013161]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.7-1.png", + "caption": "Fig. 18.7 Two pulleys connected by an elastic belt and transmitting a torque M", + "texts": [ + "2 The pointC cannot pass beyond the segment AB, so there is a limit to the tangential force that can be transmitted, equal to f P , where P is the total normal force. If this is reached the system transitions to gross slip. The cone then skids along the bar in the case of braking, or spins with uncorrelated translational motion in the case of acceleration. Johnson (1985) showed that many of the qualitative features of the effect of material deformation on rolling contact can be exposed in the simple belt drive of Fig. 18.7. Two identical rigid pulleys of radius R are connected by an elastic belt whose axial strain e is given by e = T k , (18.7) where T is the belt tension and k is a constant. If the system is running at constant speed, the pulleys must be in equilibrium and hence the driving and driven torques must both be given by M = (T1 \u2212 T2)R , (18.8) where T1, T2 are the tensions in the upper and lower free segments of belt respectively. These tensions must differ if power is to be transmitted, so the corresponding axial strains e1, e2 in the non-contacting belt segments must also differ" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003792_b978-0-12-812667-7.00031-8-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003792_b978-0-12-812667-7.00031-8-Figure6-1.png", + "caption": "FIG. 6\u2014CONT\u2019D", + "texts": [ + " Medical robotics for minimally invasive soft-tissue surgery faces significant challenges [96\u2013100] including a means for precision actuation [97, 104]. Prior work in the literature on multisteerable catheters and ultrasound-guided needles showed they are too large or lack controllable access [98\u2013100]. These limitations indicate our millirobot small size and mobility in tissue are unique. An earlier effort [99] did not produce as small of a robot as we propose partly because the 3-D microprinter leveraging technology was not available then. Small size and remote operation of the millirobot, Fig. 6A, lead to two fundamental challenges and basic science impediments, which are (i) power delivery and (ii) mobility. Power is needed for propulsion and robot functionality but has been limited because electric motors are too large and need a gearbox to produce sufficient torque, and batteries (too large) and chemical power (too low) are not practical. Mobility is critical to enable (A) (B) (C) (D) (E) (F) P C W W C D D W W P ul l w ire Tu be C N T w ire FIG. 6 Millirobot design: (A) concept snakelike millirobot (B) that can move through tissue to a tumor site and deliver medicationor performsurgery; (C)\u2013(E)multilumencoil-reinforced variable durometer tetherwith lumen for pull wires (W), drug/nanoparticle delivery (D), pressure (P), and CNT wire (C); (F) robot body side-view cross section showing CNT wire (and coil circles) and pull wire in tube. Conventionally, only 1% of any nanoparticle (NP) dose gets into tumors. In the bloodstream, proteins bind weakly to the NP and form a cloud around it, covering the tumor-binding molecules", + " (G) and (I) From Nanoscribe Inc. https://www.nanoscribe.de/en/media-press/press-releases/high-precision-micro-components- fabricated-additive-manufacturing/. 842 CHAPTER 31 CARBON NANOTUBE WIRE controlled movement through tissue, accounting for the natural heterogeneity of the in vivo environment, and to position the millirobot at target tissues but has been limited by the difficulty of moving/ steering without macerating tissue. These issues can be addressed through engineering a novel tether (Fig. 6C\u2013E) that incorporates four miniaturized Dyneema pull wires (W), hydraulic pressure (P), drug delivery/biopsy extraction tubes (D), and CNT electric wire (C). The 3-D microprinter will print novel multifunctional tips with microactuators for surgery, thermal deposition/tissue ablation, diagnostic 8436 SUMMARY sensing, and material delivery for therapy. Steering is achieved by imbalancing stress in the tube wall using strong pull cables (braided flexible Dyneema) in the lumen (W in Fig. 6C) attached to the tip. Fluoroscopy imaging for navigation provides submillimeter resolution. Millirobot Manufacturing. Novel 50\u03bcm-diameter flexible hybrid CNT-Ag wires [106\u2013109] will be manufactured in-house to power (DC or RF) the millirobot. Coiled CNT wires will generate an electromagnetic field (Fig. 6F) for heating tumor tissue or delivering or collecting magnetic nanoparticles for tumor therapy. The robot tip with CNT coils [107\u2013111] (Fig. 6F) will be 3-D printed (Fig. 6G) and metallized. The tether (Fig. 6C and E) could use high-pressure fluid to safely vibrate the robot tip to shear (microslide) through tissue. Conventional tubing is stiffer in the extruded axial direction and soft circumferentially and thus unsuitable for use as a tether. The concept tether uses a spiral (compression spring) wire-reinforced multilumen variable durometer tube enabling oscillating high pressure (water or perfluorocarbon) to axially vibrate the tip to allow a robot for the first time to go around and between organs (e", + " Controlling the movement of a tether that drives a millirobot will revolutionize multiple current medical devices including robotic arms/catheters, and it will also introduce a platform to build an extensive set of minisurgical tools to improve health care and reduce costs [112\u2013119]. Also, microscale tube fittings can be 3-D printed for use for microlumen tubing that currently requires larger Tuohy-Borst fittings. Qosina Corp. or Microlumen Co. might commercialize microfittings that would be developed for use on the millirobot. A large robot (6mm diam) moved a robot tip with a blade through simulated tissue by hydraulic pressure producing straight large vibration (7mm) of the tip, as shown in Fig. 6H. Possible applications are lung cancer screening, in vivo sensing subarachnoid space chemicals, pressure, flow, glaucoma, tubes, needles, surgery, smart implants, communication, and control. This chapter described how CNT wire can be an enabling technology for precision medical device development. CNT wire is inert, soft and lightweight. It has good conductivity, has small diameter, and good strength. These characteristics makes CNT wire an ideal material for in vivo applications. The long-term safety of CNT wire needs to be further investigated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003738_978-3-030-04975-1_7-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003738_978-3-030-04975-1_7-Figure3-1.png", + "caption": "Fig. 3. Manufacturing tolerances for the main elements of the transmission", + "texts": [ + " By inserting expressions (2), (3) (4), (5) into expression (1) and after their transformation, the total power lost in the hipocycloidal gearing is determined: NCk r\u00f0hipo\u00de \u00bc xh e rw1 Xn i\u00bc1 Fi fk r ui 2 g \u00fe fr o ui 2 g 1 \u00f06\u00de The dependencies described above relate to an ideal gearbox, made without backlash. However, its elements are made with some assumed tolerance. Thus, intertooth clearances are assumed for a gear transmission, even if to allow its assembly. The determination of the resulting clearance in mesh Di is necessary to determine the actual distance ui and the value of intertooth Fi. The manufacturing tolerances of the individual elements influencing the creation of intertooth clearance Di are shown in Fig. 3. Therefore, the following manufacturing tolerances can be distinguished: \u2022 tolerance for the gear outline - Tzh; \u2022 tolerance for the roller - Tr; \u2022 tolerance for the rollers arrangement radius - TRg; \u2022 tolerance for the angular arrangement of the rollers - TuR; \u2022 tolerance for the eccentric - Te. Based on Fig. 4, the following can be specified: \u2013 distance ui ui \u00bc rw2 yozri cos# \u00fewi \u00f07\u00de \u2013 distance wi wi \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xzh xozri\u00f0 \u00de2 \u00fe yozri yzh\u00f0 \u00de2 q \u00f08\u00de The real outline of the hypocycloidal gearing, with deviation dzh from the theoretical outline can be described in relation to the X02Y system as follows: xzh \u00bc q z 1\u00f0 \u00de cos g\u00fe k q cos z 1\u00f0 \u00deg\u00fe g\u00fe dzh\u00f0 \u00de cos g k cos z 1\u00f0 \u00degffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 k cos zg\u00fe k2 p \u00f09\u00de yzh \u00bc q z 1\u00f0 \u00de sin g k q sin z 1\u00f0 \u00deg\u00fe g\u00fe dzh\u00f0 \u00de sin g\u00fe k cos z 1\u00f0 \u00degffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 k cos zg\u00fe k2 p where: dzh \u2013 deviation of the theoretical hypocycloidal outline", + " The coordinates of the position of the roller\u2019s centre for both gearings are determined by the expression: xozri \u00bc Rg \u00fe dRg sin aki \u00fe daki\u00f0 \u00de yozri \u00bc Rg \u00fe dRg cos aki \u00fe daki\u00f0 \u00de e\u00fe de\u00f0 \u00de \u00f010\u00de where: dRg \u2013 deviation of the rollers\u2019 placement radius Rg, daki \u2013 deviation of angle aki of the rollers placement, de \u2013 deviation of eccentric e. Intertooth clearance Di (backlash) for both gearings is determined from the dependence: Di \u00bc wi \u00f0g\u00fe dg\u00de \u00f011\u00de where: dg \u2013 deviation of the roller radius The determination of intertooth clearance Di, especially its boundary values, taking into account all tolerances (Fig. 3), requires the use of probabilistic methods to determine the distribution function of that clearance. It is also possible to narrow down the search field for the value of that clearance by selecting transmission elements, e.g. a roller featuring the same deviation and a drive shaft featuring the same deviation of the eccentric. During the operation of the cycloidal gear transmission, MK torque is generated and it acts on the planetary gear. This torque generates intertooth forces Fi which occur between the radius g rollers and the hypocycloidal curve that gets in contact with them" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000622_icinfa.2009.5205077-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000622_icinfa.2009.5205077-Figure4-1.png", + "caption": "Fig. 4. Simplified robot-terrain interaction diagram.", + "texts": [ + " Thus, this method can be applied to derive more exact localization relative to terrain features that the robot must navigate, and provides localization when other traditional landmarks are unavailable. This method can eventually be combined with other traditional techniques to improve the accuracy of the robot localization in general outdoor environments. A. Terrain Map Extraction The analysis throughout this paper is based on two-axle mobile robots which are simplified as a line connecting the two wheel contact points 1C and 2C , Fig. 4. The assumption is made that only the slope at 2C can be measured. Given a path, xd, on a topographical map, a simple example of a cross-section of the topographical map { ( )}Mz z f x is shown in Fig. 5. Then a terrain inclination map { ( )}M M M dh x needs to be extracted as a reference for slope measurement at the point 2C . The variable dx represents the location of the robot on the path. In this example, the whole map is divided into eight sections. The slope function is 0( ) 0M d Mh x at 0 1b x b L or 0 0 1 1dd b x b L d where 0M is the slope of the first surface, S0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001931_etechnxt.2018.8385373-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001931_etechnxt.2018.8385373-Figure2-1.png", + "caption": "Fig. 2. Schematic for the measurement of stator phase winding resistance.", + "texts": [ + " For the calculation of the flux characteristics, the initial step is to measure the stator phase winding resistance of the motor. The stator phase winding resistance of the low voltage SRM is in general very low. For accurate resistance measurement, voltmeter-ammeter technique is used. Due to the low resistance of the stator phase windings, a small voltage is applied at the terminals of the winding to maintain the current within limit. For higher voltage, a resistance Rx in series is connected. In Fig. 2, the schematic for the estimation of stator phase winding resistance (R) is shown. The winding resistance is calculated from Vw/Isw. The mean value of resistance R, is \u03a8 calculated from the measurement and it is achieved as 0.143\u2126. For determining the flux linkage, voltage across and current through across the SRM phase windings, are measured. The current measurement starts from zero until the steady state value is achieved. Normally it is 200% of the rated value. Based upon the stator phase winding resistance of the SRM, the battery voltage has to be calculated, which is to be applied across the SRM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003110_978-981-13-2375-1_67-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003110_978-981-13-2375-1_67-Figure2-1.png", + "caption": "Fig. 2. The schematic of the 3-PRS platform", + "texts": [ + " 3, the model of identification is deducted by transforming the dynamic model into the linear equation about the parameters to be identified and the stepwise identification method is proposed to identify dynamic parameters. In Sect. 4, the effectiveness of parameters identification methods are verified. Some conclusions are drown in Sect. 5. The structure of the flight simulator motion platform studied in this paper is illustrated in Fig. 1. It has three degrees of freedom including pitch, roll and vertical movement. The schematic of the simulator platform is shown in Fig. 2. The intersections of the base and stand columns are represented by Ai\u00f0i \u00bc 1; 2; 3\u00de. The point O represents the center of the fixed base, and OAik k \u00bc a i \u00bc 1; 2; 3\u00f0 \u00de. The ends of limbs and centers of sliders are crossing at Bi\u00f0i \u00bc 1; 2; 3\u00de, and AiBik k \u00bc Pi. The connections of ends of limbs and moving platform which are also centers of ball hinges are represented by Ci(i \u00bc 1,2,3). The point O1 represents the center of moving platform, and OCik k \u00bc b i \u00bc 1,2,3\u00f0 \u00de. The length of limb is BiCik k = l. In order to analyze the kinematics of the parallel platform, a fixed reference system O XYZ and moving frame O1 uvw are established, shown in Fig. 2. According to the structural characteristics, the rotations around the X and Y axis are the two rotational degrees of freedom of the platform. The Euler transformation matrix R [9] of O1 uvw related to O XYZ is obtained by the \u201cZ-X-Z\u201d Euler transformation method. The Euler angle of c is equal to a. Then the coordinates of Ci and Bi in the fixed reference system can be computed. The closed-vector method is used to compute the displacement of all moving parts, and other kinematic parameters including velocity, acceleration and angular velocity can be derived" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001617_s40799-018-0232-7-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001617_s40799-018-0232-7-Figure6-1.png", + "caption": "Fig. 6 Schematic of connection", + "texts": [ + " 5, where a frame that holds a circular hollow bar, a solid cylinder and a calibrated mass is observed. A strain gage, type T, is placed on the solid cylinder. This gage measures strains in orthogonal directions simultaneously, exactly as shown in the numerical model. However, only one of these directions is comparable with the proposed analytical model. The signal is sent from the strain gage to an analogue data acquisition card, National Instrument\u00ae 9219, which is connected to a USB carrier NI 9171 that sends the signal to a computer through a USB cable, as shown in Fig. 6. Each signal is sent to two independent channels, each in the configuration of a quarter bridge, as shown in Table 3. The data received by the computer are recorded and processed in a virtual instrument developed in the LabVIEW\u00ae Academic Suite software. The strain gage used in this work is a CEA-06-125UT-350, type T with grids at 90\u00b0 for general purposes. These grids, 1 and 2 (see Fig. 7), have slightly different gage factors of 2.165 and 2.185, as well as transverse sensitivity coefficients of 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003675_022092-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003675_022092-Figure4-1.png", + "caption": "Figure 4 Load applied", + "texts": [ + " The specific boundary conditions are applied as shown in Table 1 and Figure 3. According to the real situation of the research object and research needs, the load of force and displacement load conditions of the model are set in this study. Concentrated force is applied to the reference point corresponding to the inner ring. This setting uniformly applies the concentrated force load to the inner surface of the inner ring of the bearing so as to avoid excessive local force caused by direct application of the concentrated force. Figure 4 shows load application, where the longer yellow arrow whose starting point is the origin of the coordinate represents the direction of the gravity load, while the shorter yellow arrow whose starting point is the reference point (RP) represents the direction of the concentrated force with the value of 17009 N. CISAT 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1168 (2019) 022092 IOP Publishing doi:10.1088/1742-6596/1168/2/022092 The Euler body is used to represent the distribution of volume fraction and the state of motion of the fluid, including the lubricating oil" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000406_imece2008-67991-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000406_imece2008-67991-Figure1-1.png", + "caption": "Figure 1. Schematic of the LSFF process", + "texts": [ + " In Laser Solid Freeform Fabrication (LSFF), a fully functional component can be fabricated directly from its CAD model by melting and subsequent solidification of additive materials injected into a melt pool formed by a laser beam on a moving substrate. Compared to other comparable conventional techniques, LSFF has shown many technical and manufacturing process planning advantages such as a small heat affected zone (HAZ), minimal dilution, and specific size raw material procurement. Hence, LSFF has been recognized as a potential candidate for different manufacturing applications including rapid manufacturing, production of functionality graded parts, and part repair, especially for high-value tools [ 1 - 4]. Figure 1 shows a schematic of the LSFF process using powder injection system for the material deposition in which a laser processing head, a lateral nozzle, and a substrate are shown. In spite of all demonstrated and inherent advantages of the LSFF process, the parts fabricated using this technique are prone to a number of drawbacks namely porosity, delamination, 1 Copyright \u00a9 2008 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use Dow and crack formation which are mainly prevalent in layered manufacturing [ 5, 6]", + ", uz=0 which is the velocity is the z direction), the initial-essential and the optimum boundary conditions for thermal governing equation are derived as: 0 ( , , , ) when 0 andT x y z t T t t= = = \u221e (3) 2 2 2 2 2[( ) ( ) ]2 ( ) | exp x y e l l l x u t y u t K T n P r r \u211c \u2212 \u2212 + \u2212 \u2212 \u2207 \u22c5 = \u03b2 \u03c0 0 ( ) if c h T T\u2212 \u2212 \u211c \u2208 \u039b 0 ( ) if c h T T= \u2212 \u2212 \u211c \u2209 \u039b (4) where \u03b2e is effective absorption factor, Pl (W) is the total laser power, ux (m/s) and uy (m/s) are the laser beam velocities in the x and y directions respectively, rl (m) is the laser beam radius, hc (W/m 2 .K) is combined heat transfer coefficient, T0 (K) is the ambient temperature, \u211c (m 2 ) is the substrate surface, and \u039b (m 2 ) is the area of the laser beam on the substrate. \u211c and \u039b are shown in Figure 1. For stress and strain fields, initial conditions and the stresses in terms of the strains represented by the thermoelastic constitutive equations, Equation (2), can also be expressed as [ 26]: ( , , ,0) | 0 & ( , , ,0) | 0 o oT T T T x y z x y z = = \u03b5 = \u03c3 = (5) [ (1 2 ) (1 )(1 2 ) ij ij kk ij E \u03c3 = \u03bd\u03b4 \u03b5 + \u2212 \u03bd \u03b5 + \u03bd \u2212 \u03bd (1 ) ] ( , , 1, 2,3) ij T i j k\u2212 + \u03bd \u03b1 \u2206 \u03b4 = (6) where \u03bd is Poisson\u2019s ratio, \u03b1 (m/m.K) is the linear coefficient of thermal expansion and \u03b4ij is the Kronecker delta. 3 Copyright \u00a9 2008 by ASME ms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-1-1.png", + "caption": "Fig. 19-1: Motor cross section Fig. 19-2: Independently excited motor", + "texts": [ + " In these applications mostly a PM-motor with chopper is used. http://dx.doi.org/10.1016/B978-0-12-814641-5.50004-4, Copyright \u00a9 2018 Elsevier Inc. All rights reserved. Power Electronics: DRIVE TECHNOLOGY and MOTION CONROL 19.2 SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR A. DC-MOTOR supplied from an AC POWER GRID 1. CONTROL OF AN INDEPENDENTLY EXCITED MOTOR 1.1 Generalities A classic DC motor is constructed from a stator with field winding F 1 F 2 and a rotor with commutator, brushes and an armature winding A 1 A 2 : see fig. 19-1. The armature winding sits in the groves of the laminated rotor. The field winding F 1 F 2 may also be replaced by permanent magnets. To counteract the armature action machines have auxiliary poles n h - z h (for power levels from 1kW) and a compensation winding C 1 C 2 (above 100 kW). This compensation winding sits in the teeth of the main poles and is connected in series with the armature winding. The flux \u03a6 in the machine is proportional with the excitation current I m and depends upon the magnetic induction and the cross sectional area of the iron" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure2-1.png", + "caption": "FIGURE 2 Gear-shaft interference assembly FEA model", + "texts": [], + "surrounding_texts": [ + "\u00a9 2018 SAE International. All Rights Reserved.\nAutomotive timing gear trains, transmission gearboxes, and wind turbine gearboxes are some of the application examples known to use interference-fit to attach the gear to the rotating shaft. This paper discusses the interference-fit joint design and the finite element analysis to demonstrate the distortion. The mechanism of tooth profile\ndistortion due to the interference-fit assembly in gear trains is discussed by demonstrating the before and after assembly gear profile measurements. An algorithm to calculate the profile slope deviation change is presented. The effectiveness of the computational algorithm to predict the distortion is demonstrated by comparing with measurements. Finally, steps to mitigate the interference assembly effects are discussed.\nGearboxes are used to transfer rotational motion between two shafts by the means of gears. The gears and shafts are designed such that there is no relative motion between them. Any relative motion can result in undesired consequences including failure. In engine timing gear train that are used to transfer power from the crank shaft to the cam shaft, accessory shafts, and power take-offs, accurate timing is maintained in the relative positions, as catastrophic damage can result due to the relative motion between the shaft and the gear. The gear and shaft joints can be of the following types:\n1. Integral 2. Bolted 3. Spline 4. Key 5. Taper 6. Interference-fit\nInterference-fit joints are designed, as the name says, to result in interference between the shaft and the gear. The resultant joint is a simple way of connecting a gear to a shaft. Such an arrangement is used commonly to connect a crank gear to a crank shaft. It is also sometimes used to connect gears to shaft in transmissions, and sun and planet gears to shaft and housing in wind turbine and other applications. An interference-fit connection is sometimes used to connect a ring gear to a housing in planetary gearboxes. The effect of the interference-fit joint design on gear profile has been reported by the author previously,\nbut a detailed understanding of the phenomenon and the required design changes to mitigate the effects is required.\nInterference-fit joints work by producing pressure between the outer member, gear hub, and the inner member, shaft by designing a difference in the diameters. The contact pressure is the result of the interference. The applied torque is reacted by a friction moment, which results from the friction and contact pressure between the contact surfaces. Gears are usually assembled to the shaft/ housing either by heating/ cooling the gear and then shrinking it on the shaft/ housing or by pressing it in position over the shaft/ housing. The interference-fit joint is characterized by radial and hoop stresses in the gear body [3]. The interference-fit joint analysis can be done accurately by Finite Element Method [4]. Qiu et al. discuss an iterative analytical method for\u00a0 calculating results of interference-fit joints with good accuracy.\nThe transmission error, which is the difference in the ideal position and the actual position of a driven gear along the plane of action, is the result of gear profile deviation. Houser, Blankenship [5, 6, 7] have published work that correlates the transmission error with the gear whine noise. The distortion can also be a result of temperature increase during operation, as well as deflection due to mesh load [8]. Drago et al. discuss the effect of the rim distortion on the gear stresses [9]. Kashyap et al. discuss the tooth geometry change of plastic gears due to temperature increase [10]. Joshi discusses the gear whine noise due to interference-fit gears. The mechanism and the effect of interference-fit joint on the gear tooth distortion and the resultant increase in profile", + "slope deviation and transmission error has not been discussed previously in the literature.\nThe interference-fit (IF) joint typically consists of a gear body with teeth and rim assembled on a shaft and web and hub may or may not be present. The IF joint is an economical way of connecting the gear body to the shaft without the use of spline, bolts, or keys. It allows connection of a gear with a shaft of comparable diameter where there is not enough material to design a spline connection at low cost. Due to this inherent motivation for the IF joint, the gear rim thickness is an important design consideration to avoid thin rim effects such as- effect of rim thickness on root stress, effect of rim thickness on crack propagation, and the effect of rim thickness on mesh stiffness. The IF gear is usually designed to have a back-up ratio i.e. the tooth whole depth to the rim thickness ratio greater than 1.2. A typical gear design has the back-up ratio value between 1.2 and 2.\nFigure 1 shows an interference-fit gear on a shaft. Table\u00a01 shows the geometry details of the gear and the shaft for the reference design discussed in detail and henceforth called design A.\nThe gear-shaft joint in Figure 1 was analyzed by using Finite Element method.\nThe gear-shaft joint was analyzed using a Frictional contact type joint in ANSYS Workbench software. As the gear-shaft joint and the effect on gear are the interest areas of the analysis, only the portion of the shaft close to the gear mounting was included in the finite element model. The shaft was fixed at the centerline at one end and there was no other force acting on the model. The static structural analysis was carried out for the gear-shaft interference set at nominal, minimum, and maximum interference. X axis was oriented radially, Y axis denoted angular orientation, and Z axis was aligned along the centerline of the gear. FEA results are listed in Table 2.\nFigures 3 and 4 show the radial deformation along the X axis in the gear body for nominal interference. The deformation due to interference is maximum at near the interface and reduces slightly with increase in diameter. The maximum radial deformation increases with increase in interference.\nFigure 3 (center) shows the circumferential deformation in the gear body for nominal interference. The circumferential deformation due to interference is small and varies slightly from one end to the other. The maximum deformation increases with increase in interference.", + "\u00a9 2018 SAE International. All Rights Reserved.\nFigure 3 (right) shows the axial deformation in the gear body for nominal interference. The axial deformation due to interference is small. The maximum deformation shows slight increase with increase in interference.\nOverall, the gear body shows a significant radial increase in diameter with minimal changes in circumferential and axial directions." + ] + }, + { + "image_filename": "designv11_92_0003145_032005-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003145_032005-Figure4-1.png", + "caption": "Figure 4. Idealized arc-shaped traction model.", + "texts": [ + " Specifically, the maximum traction force required for laying high-voltage cables in different lengths and different types of stay tubes can also be calculated, with reference to the laying trajectory model given in the actual laying scheme. Consistent with what has mentioned previously, a simplified diagram of the relevant model will be given at first. In order to obtain a simplified model of the arc-shaped vertical laying under high-drop section, the High Drop Section portion in Fig. 1 is simplified to the model in Fig. 4 below. According to the force analysis, the required traction force can be obtained for each segment: Force in arc DC segment: IMMAEE 2018 IOP Conf. Series: Materials Science and Engineering452 (2018) 032005 IOP Publishing doi:10.1088/1757-899X/452/3/032005 2 2 9.8 [2 sin (1 )( cos )] 1c D WR T e T e (7) Force in CB segment: 9.8 2 ( cos sin ) 4 4B c BCT T W L (8) Force in arc BA segment: 2 2 9.8 [(1 ) sin 2 ( cos )] 1A B W R T T e e (9) The maximum allowable traction of the cable will be further considered: mT A (10) In the above formula, is the allowable traction strength of the conductor(N/mm2) and A is the cable conductor cross-sectional area(mm2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002103_ccdc.2018.8408202-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002103_ccdc.2018.8408202-Figure1-1.png", + "caption": "Fig 1. The relationship of Leader-Follower UAVs", + "texts": [ + " Therefore, the state equation of the system can be transformed into i i i i p q q u = = (7) After feedback linearization, the new input i\u03b7 is obtained. In formula (7), ip and iq represent the position locations and the velocity magnitudes, respectively. Then, it can be written as i i iX AX Bu= + (8) where, i i i p X q = , 3 3 3 3 3 3 3 3 0 0 0 I A \u00d7 \u00d7 \u00d7 \u00d7 = , 3 3 3 3 0 B I \u00d7 \u00d7 = . 2.2 Three-dimensional Elastic Distance Vector The geometrical relationship between the leader and the follower UAV in the formation is shown in Fig. 1. 0P and iP represent the mass center of the leader vehicle and the ith follower in the inertial reference frame, respectively. iP and i\u0302P represent the desired look-ahead point and the elastic look-ahead point of ith UAV, respectively. i i i iPx y z denotes the aircraft-body coordinate frame of the ith UAV. g g g gO x y z denotes the ground coordinate system. The conversion relationship between the aircraft-body and the ground coordinate frame is i i i p X q = , 3 3 3 3 3 3 3 3 0 0 0 I A \u00d7 \u00d7 \u00d7 \u00d7 = , 3 3 3 3 0 B I \u00d7 \u00d7 = (9) g id and p id represent the distance vector in the ground and the aircraft-body coordinate frame, respectively, and satisfy the following relationship g T p i id S d\u03c3\u03d5= (10) The elastic distance vector \u02c6 g id is defined as ( ) ( ) \u02c6 / g g i i i i i d d f v d f v v v = = (11) where, dv is a constant, denotes the expected flight speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002826_detc2018-86294-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002826_detc2018-86294-Figure4-1.png", + "caption": "Figure 4 Finite element anlysis model of gear transmission system", + "texts": [ + " The Matern covariance function and Gaussian likelihood function are chosen in this study. The results based on GPR and MC simulation are displayed in Table 3. Due to the stochastic nature of the environmental loads and time-varying internal incentive, it is challenging to include uncertainties for the assessment of the fatigue damage accumulations in the gear transmission system\u2019s life cycle. In this paper, the probabilistic fatigue is estimated using finite element analysis method and GPR method. The finite element analysis model is provided in Figure 4. 4 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 11/18/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use In this case study, firstly, the dynamic contact stress in a meshing period under different working conditions are analyzed. Secondly, for each working condition, the distribution of the contact stress in a meshing period is calculated, which is lognormal distribution. The distribution information is listed in Table 4. The data of working condition 1 to working condition 10 are defined as training data, the rotating speed of 140rad/s and the torque of 130N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000815_icelmach.2008.4800217-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000815_icelmach.2008.4800217-Figure3-1.png", + "caption": "Fig. 3. Schematic diagram of the magnetic circuit at the stator pole.", + "texts": [ + " These are given from the number of winding turns N and each phase current iu, iv, and iw, respectively. The Rsps, which are nonlinear reluctances on each stator poles, are given by following equation: \u239f \u239f \u23a0 \u239e \u239c \u239c \u239d \u239b += 12 13 131 \u03c6 \u03b2\u03b2 sp sp sp sp sp S l S l R (1) lR uNi wuR uvR lR lR lR lR lR vwR uvR wuR vwR spR uNi wNi vNi vNi wNi mu\u03c6 mv\u03c6 mw\u03c6 mu\u03c6 mw\u03c6 mv\u03c6 spR spR spR spR spR Fig. 2. Magnetic circuit model of the IPM motor. 978-1-4244-1736-0/08/$25.00 \u00a92008 IEEE 1 The schematic diagram of the magnetic circuit at the stator pole is illustrated in Fig. 3. Ssp is cross-sectional area of the stator pole and lsp is the length of the stator pole. \u03b21 and \u03b213 are the first order and the 13th order magnetization coefficient of the material, respectively. \u03c6mu, \u03c6mv, and \u03c6mw are the each phase fluxes from the permanent magnets, respectively. The reluctances between each phase are Ruv, Rvw, and Rwu, respectively. These reluctances and magnet fluxes change with the rotor angle. In order to obtain the reluctances and magnet fluxes, each flux passing through the stator pole are calculated using finite element method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002708_978-3-319-94476-0_5-Figure5.15-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002708_978-3-319-94476-0_5-Figure5.15-1.png", + "caption": "Fig. 5.15 Illustration of two geometries implemented in COMSOL, together with the mesh generated for one of the geometries. The samples represent 2D rectangles of aluminum with cracks (horizontal and inclined at 20\u25e6) of finite length positioned in the center domain. Both geometries are meshed with triangular mesh elements, with a higher mesh density in the region of the crack", + "texts": [ + " In order to illustrate the potential of the proposed model, two instructive examples of a shear wave propagating in a 2D rectangular aluminum sample of 50 mm width and 100 mm height containing a crack are studied. The aluminum sample has a density \u03c1 = 2700 kg/m3, Young\u2019s modulus E = 70 GPa, and Poisson\u2019s ratio \u03bd = 0.33. A crack with a length of 20 mm is positioned in the center of the sample. In the first example, the crack orientation is horizontal, whereas in the second example the crack is inclined at 20\u25e6, as illustrated in Fig. 5.15. In both examples, a continuous shear wave excitation with a frequency of 100 kHz and tangential displacement amplitude of 100 nm is defined on the top boundary of the sample. At the side and bottom boundaries of the sample, non-reflecting boundary conditions were applied in order not to mask the crack-wave interactions by parasite reflections. At the internal crack boundaries, a thin elastic layer boundary condition is specified as described in Sect. 5.8.2. The friction coefficient value \u03bc = 1 was used, which is close to known data for aluminum on aluminum [38]. As illustrated in Fig. 5.15, the full geometry is meshed using triangular elements with a maximum size of approximately 2.6 mm (i.e., 12 second-order mesh elements per wavelength). At the internal crack boundary, however, a fixed number of 150 mesh elements (i.e., element size of approximately 0.13 mm) was used. By choosing such small elements, the macroscopic elastic fields used in the MMD algorithm can be considered uniform enough within each mesh cell. The solution to the problem is calculated using the implicit generalized alpha time-dependent solver typically used for structural mechanics problems in COMSOL" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001951_icaset.2018.8376937-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001951_icaset.2018.8376937-Figure1-1.png", + "caption": "Fig 1: experiment setup of the system.", + "texts": [ + " The high-gain observer [5] has excellent properties in states observation for a large gains, the main difficulty in this method is determining the suitable gains. Luenberger observer [6] depends on system\u2019s dynamics model and measured inputs and outputs to observe system states along with the disturbance, as this method completely depends on the dynamic model, any change in system\u2019s dynamics, will cause the observation error to become larger. System description The system consists of a servo motor connected to an elastic element through a rack and pinion, with a mass attached on the other side of the elastic element as shown in Fig.1. When the shaft of the motor rotates, the angular motion is converted into a linear motion, which affect on the spring by a force opposite to spring force due to Hook\u2019s law, so the mass moves. The angular position of the shaft can be easily measured using an encoder; therefore, rack pinion position can be obtained by the relation . Where x is the position of the rack, r is the x r radius for the pinion, and is angular position of motor\u2019s shaft. However, the position of the mass is assumed to be unmeasured directly through a sensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003173_012129-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003173_012129-Figure1-1.png", + "caption": "Figure 1 Working Principle of Laser Cladding", + "texts": [ + " These components erode, corrode and are subjected to severe wear. It will be very expensive to replace these components. In order to overcome these problems of replacement, laser cladding is one such technique where a thick coating can be applied on the surface subjected to wear and tear with the required metal powders. This process gives a better coating with minimum dilution, minimal distortion, least HAZ, better surface quality and also requires reduced post machining as compared to other conventional coating processes of spraying [1]. Figure 1 gives the basic principal of laser cladding process [2]. The temperature variation during the process of cladding is an important parameter to be observed, because of the fact that the temperature of the substrate beyond a certain limit may lead to distortion of the workpiece as similar to welding process. This distortion even at micron level is not allowed in components like aerospace application, etc. R. Jendrzejewski et al. studied the distribution of temperature in multi-layer laser cladding using numerical analysis approach and thereby compared the IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000551_icara.2000.4803997-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000551_icara.2000.4803997-Figure4-1.png", + "caption": "Fig. 4. Three dimensional probability clouds of two indoor GPS measurements in three dimensional state space. The height of these clouds represent the measured angle, the diameter in height is the variance of the angle \u03c3\u03d5. The same holds for the measured position in x and y with the corresponding variances \u03c3x and \u03c3y as diameters.", + "texts": [ + " As example for the inaccuracy the error in the x-component is shown in Tab. I. What remains is a large statistical error that is increasing from the middle to the sides of the room (for x see Tab. I). To generate a probability grid of one indoor GPS measurement, this inaccuracy has to be considered. We assume a Gaussian distributed statistical error in all three state space directions. An example of two measurements and its probability distribution clouds in the state space (x, y, \u03d5) is shown in Fig. 4. Their sizes differ because of the different variances. As mentioned the variances \u03c3x, \u03c3y and \u03c3\u03d5 of all dimensions increase to the sides of the mission space, the larger the variance the larger is the diameter of the cloud in this direction. The output of the calibrated indoor GPS system is one cloud (one measurement) in the three dimensional state space as position probability grid Mgps(x, y, \u03d5). Fig. 5 shows a two dimensional Gaussian distribution in a cross section at an angle \u03d5 of the resulting three dimensional grid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001479_978-3-319-72730-1_26-Figure26.10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001479_978-3-319-72730-1_26-Figure26.10-1.png", + "caption": "Fig. 26.10 Root loci with saturation for voltage-fed excitation", + "texts": [ + " Introducing three additional parameters \u03c4 f = l f t/r f , \u03c3 f = l\u03c3/ l f t and \u03bb = maf / l f t , the characteristic equation can be written as follows: 1 + (\u03c4ai/\u03c4mi ) { 1 + p\u03c4 f [\u03c3 f + (\u03c4ai\u03c4mi/\u03c4a\u03c4m)(1 \u2212 \u03c3 f )] } [ (1 + p\u03c4ai )(1 + p\u03c4 f ) + p\u03c4 f \u03c4ai (1 \u2212 \u03c3 f )\u03ba(\u03bdro \u2212 \u03bbp) ] (\u03c4ai/\u03c4w + p\u03c4ai ) = 0 (26.59) There is now a finite zero around p \u2248 \u2212\u03c4\u22121 f . The open-loop poles include the mechanical damping pole and a pole around \u2212\u03c4\u22121 a as well as a third open-loop pole around the same value as the finite zero, i.e. p \u2248 \u2212\u03c4\u22121 f (without saturation and armature reaction, this pole and this zero coincide so that the system then becomes second order). Possible root loci for motoring and generating are depicted in Fig. 26.10 ((a) for motoring and (b) for generating). As is the case for a current-fed excitation, the saturation and armature reaction stabilise the behaviour for generating, while for motoring the damping (of the dominant eigenvalue) decreases." + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001813_978-981-10-8306-8_13-Figure13.7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001813_978-981-10-8306-8_13-Figure13.7-1.png", + "caption": "Fig. 13.7 Multiple plastic strain increments a by the simple tension and b balanced biaxial stress states", + "texts": [ + " However, there are multiple stress states for plane strain deformation from the simple tension to the balanced biaxial. For the pure shear stress state, deformation of the Tresca case and all incompressible, isotropic and symmetric yield functions including the von Mises case is the same as plane strain deformation, even though the stress state for the plane strain deformation is non-unique for the Tresca case. The multiple plastic strain increments for the simple tension are depST \u00bc depSTI 0 0 0 \u00f0 1\u00fe c\u00dedepSTI 0 0 0 cdepSTI 0 B@ 1 CA \u00f013:36\u00de where 0:0 c 1:0 as shown in Fig. 13.7a. For c \u00bc 0:5, deformation becomes equivalent with that of simple tension for the smooth incompressible and isotropic yield function. For c \u00bc 0:0 and c \u00bc 1:0, the deformation is the plane strain deformation, which is also obtainable with the pure shear stress state as shown in Fig. 13.7a. The figure confirms that depSTI \u00bc depPSI \u00bc dY \u00bc 1 2 dK \u00f013:37\u00de which complies with Y = 2K, as confirmed by the plastic work equivalence principle. The multiple plastic strain increments for the balanced biaxial stress state are depBB \u00bc \u00f01 c\u00de depBBIII 0 0 0 c depBBIII 0 0 0 depBBIII 0 B@ 1 CA \u00f013:38\u00de where 0:0 c 1:0 as shown in Fig. 13.7b. For c \u00bc 0:5, the deformation becomes equivalent with that of the balanced biaxial state for the smooth incompressible and isotropic yield function. For c \u00bc 0:0 and c \u00bc 1:0, the deformation is the plane strain deformation, which is also obtainable with the pure shear stress state as shown in Fig. 13.7a. The figure confirms that depBBIII \u00bc depPSI \u00bc dB \u00bc 1 2 dK \u00f013:39\u00de Which complies with Y = B = 2K, as confirmed by the plastic work equivalence principle. The p diagram of the plastic strain increment surface shown in Fig. 13.5b implies that the conjugate Tresca effective plastic strain increment is d e \u00bc b deIj j \u00fe deIIj j \u00fe deIIIj j\u00f0 \u00def g \u00f013:40\u00de as the p diagram of Eq. (12.35) with M = 1.0 shown in Fig. 12.20 suggests. Plot the p diagram of Eq. (13.40), which should be the same as that shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001996_icasi.2018.8394258-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001996_icasi.2018.8394258-Figure3-1.png", + "caption": "Fig. 3 Triangular pyramid CMGs configuration.", + "texts": [ + " In addition, m and [ , , ]T x y zI I II denote the total mass and inertia matrix of the CMGactuated underwater vehicle, respectively. 2 [ , , ]Tp q rv is the body-fixed angular velocity vector, and the total angular momentum vector ( )H generated by the CMG system can be expressed as 1 2 2 3 3 0 1 2 2 3 3 1 2 3 60 60 60 60 ( ) 60 60 60 60 c s s c c c s s c c c s J c c c c s s c c c s s s s s s s s H (6) where the pyramid skew angle is set as 70.53 , 1 2 3 4[ , , , ]T is the gimbal angle vector and 0J denotes the momentum magnitude for each CMG (see Fig. 3). Alternatively, the time derivative of the CMG angular momentum ( )H in Eq. 6 can be calculated as 1 2 2 3 3 0 1 2 2 3 3 1 2 3 0 60 60 60 c 60 ( ) 60 60 60 c 60 ( ) c c s s c c c s s c c J s c s c s c c s s c s c s c s c J C H where ( )C is the Jacobi matrix of the CMG system. Therefore, its singularity measure D which represents the distance from singularity is given by (7) 29ISBN 978-1-5386-4342-6 det( ( ) ( ))TD C C (8) The main focus of this paper is the attitude control of a CMG-actuated underwater vehicle at low speed, and thus we set 0u v " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000007_amr.44-46.829-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000007_amr.44-46.829-Figure5-1.png", + "caption": "Fig. 5 K5 type of titling bogie and its mechanics principle of China freight car", + "texts": [ + " But a check for flaws and inclusions indicates that they are also in reasonable range. Hardness values of the surface layer and core in bearing outer ring are checked and given in Table 2. This verifies that the hardness is obviously less than the required values. This means that the manufacturing technique may not match the technical requirements. K5 Type Bogie. An elastic-plastic finite element (EPFE) analysis has been made to the bearing on the freight car with K5 type tilting bogie [5]. As shown in Figure 5, this tilting bogie has a special characteristic that the force from side frame of the bogie is transferred to bearing through the wing block acting on the adapter. The wing block can wave right or left around its rotational centre with vehicle vibrating laterally. The wheel set is subjected to a dynamic variable balance force system, as shown in Figure 6. A wheel-track contact force spectrum inspected on-line on a R300 curved line is used for studying the bearing distributed fatigue stresses. With respect to the selected inspection positions shown in Figure 7, results of the EPFE analysis are exhibited in Table 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001042_20090622-3-uk-3004.00054-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001042_20090622-3-uk-3004.00054-Figure1-1.png", + "caption": "Fig. 1. Phase portrait of system (10) for \u00b5 = 0.5 with trajectories through tangent points and the origin. In addition, the equilibria are depicted with asterisks, and the manifolds of the saddle point are shown.", + "texts": [ + " Consider the continuous system: x\u0307 = A1x + \u00b5b, x \u2208 S1 := {x \u2208 R 2|nT 41x < 0 \u2227 nT 12x > 0}, A2x + \u00b5b, x \u2208 S2 := {x \u2208 R 2|nT 12x < 0 \u2227 nT 23x > 0}, A3x + \u00b5b, x \u2208 S3 := {x \u2208 R 2|nT 23x < 0 \u2227 nT 34x > 0}, A4x + \u00b5b, x \u2208 S4 := {x \u2208 R 2|nT 34x < 0 \u2227 nT 41x > 0}, (10) where the normal vectors are chosen as n12 = [0 1] T , n23 = 1 \u221a 2 [\u22121 \u22121] T , n34 = [0 \u22121] T , n41 = 1 \u221a 2 [1 1] T . The vector b = [cos(0.375\u03c0) sin(.375\u03c0)] T and \u00b5 \u2208 R is the bifurcation parameter. The phase portrait of this system for \u00b5 = \u22120.5 is shown in Figure 1. The matrices Ai are A1 = [ \u22120.5 1 \u22121 0 ] , A2 = [ \u22120.5 0.91 \u22121 0.58 ] , A3 = [ \u22121 0.41 0.5 2.08 ] , A4 = [ \u22121 0.5 0.5 1.5 ] . System (10) will be analysed with the given procedure: 1. For \u00b5 < 0, two equilibria exist, with positions x = \u2212\u00b5A\u22121 2 b in S2 and x = \u2212\u00b5A\u22121 4 b in S4. For \u00b5 > 0, no equilibria exist. 2. At \u00b5 = 0, the conewise linear dynamics is unstable, since the visible eigenvector in S4 corresponds to an unstable eigenvalue. In addition, one visible eigenvector in S3 exists, that corresponds to a stable eigenvalue. 3. For \u00b5 = \u22120.5: a. On \u03a312, \u03a334 and \u03a341, there exist points where the vector field is tangent to the boundary, i.e. points T12, T34 and T41, respectively. Trajectories through these points and the origin are shown in Figure 1. An unstable focus exist in S2, since the eigenvalues of A2 are 0.42\u00b1 0.79\u0131, where \u01312 = \u22121. A saddle point exist in S4 with eigenvalues \u22121.10 and 1.60. The depicted stable and unstable manifolds of this point are shown and do not form a homoclinic orbit. b. The trace tr(A1) < 0, whereas all other traces tr(Ai) > 0, i = 2, 3, 4. Therefore, application of Theorem 2 yields that each possible closed orbit visits S1. To satisfy Theorem 1, closed orbit(s) should encircle the focus without encircling the saddle point. By studying the depicted trajectories, one can conclude, that no closed orbit can traverse \u03a312 \\ [O, a], since these trajectories cannot encircle the focus without encircling the saddle point, which is required according to Theorem 1. Furthermore, closed orbits can not traverse the interior of the line [a, b], since trajectories through this open line will arrive at the line [c, d] in finite time, and enter the positively invariant region that is depicted gray in Figure 1. Now, one can conclude, that possible closed orbits visit only the regions S1 and S2, such that they should be contained in the domain, that is depicted gray. This implies, that all closed orbits traverse the line [T12, e]. c. Existing closed orbits should traverse the line [T12, e]. We construct a map g2 : [T12, e] \u2192 [d, T12], that yields the position g2(x) where a trajectory leaves the cone S2 when this cone was entered at x. Similarly, the map g1 : [b, T12] \u2192 [T12, e] is computed. The maps are computed according to (A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001937_picc.2018.8384787-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001937_picc.2018.8384787-Figure5-1.png", + "caption": "Fig. 5: Disc Discretization", + "texts": [ + " \u21d2 \u2212\u2212\u2192 \ud835\udc35\ud835\udc672 = ( \ud835\udf070\ud835\udf0f 4\ud835\udf0b ) \ud835\udc4e\u222b 0 2\ud835\udf0b\u222b 0 { [\ud835\udc3d\ud835\udc5f\u2032\ud835\udc5f sin (\ud835\udf03 \u2212 \ud835\udf03\u2032)] \u2223\u2212\u2192\ud835\udc5f \u2212\u2212\u2192 \ud835\udc5f\u2032 \u22233 } \u2212 { [\ud835\udc3d\ud835\udf03\u2032(\ud835\udc5f cos (\ud835\udf03 \u2212 \ud835\udf03\u2032)\u2212 \ud835\udc5f\u2032)] \u2223\u2212\u2192\ud835\udc5f \u2212\u2212\u2192 \ud835\udc5f\u2032 \u22233 } \ud835\udc5f\u2032\ud835\udc51\ud835\udc5f\u2032\ud835\udc51\ud835\udf03\u2032 (9) Now, define a stream function ( \u2212\u2192 \ud835\udc48 ) so that, \u2212\u2192 \ud835\udc3d = \u2207\u00d7\u2212\u2192 \ud835\udc48 (10) From Maxwell\u2019s equation, \u2207\u00d7\u2212\u2192 \ud835\udc38 = \u2212\u2202 \u2212\u2192 \ud835\udc35 \u2202\ud835\udc61 = \u2212\ud835\udc57\ud835\udf14 \u2212\u2192 \ud835\udc35 (11) where, \u2212\u2192 \ud835\udc35 = \u2212\u2212\u2192 \ud835\udc35\ud835\udc4d1 + \u2212\u2212\u2192 \ud835\udc35\ud835\udc4d2 (12) Also, \u2212\u2192 \ud835\udc3d = \ud835\udf0e \u2212\u2192 \ud835\udc38 (13) Now, by using equations (7) to (13) the stream function can be expressed as, 1 \ud835\udc5f \u2202 \u2202\ud835\udc5f (\ud835\udc5f \u2202\ud835\udc48\ud835\udc67 \u2202\ud835\udc5f ) + 1 \ud835\udc5f2 \u22022\ud835\udc48\ud835\udc67 \u2202\ud835\udf032 + \ud835\udc57\ud835\udf14\ud835\udf0e( \ud835\udf070\ud835\udf0f 4\ud835\udf0b ) \ud835\udc4e\u222b 0 2\ud835\udf0b\u222b 0 { [\ud835\udc5f cos (\ud835\udf03 \u2212 \ud835\udf03\u2032)\u2212 \ud835\udc5f\u2032] \u2223\u2212\u2192\ud835\udc5f \u2212\u2212\u2192 \ud835\udc5f\u2032 \u22233 \u2202\ud835\udc48\ud835\udc67 \u2202\ud835\udc5f\u2032 } \ud835\udc5f\u2032\ud835\udc51\ud835\udc5f\u2032\ud835\udc51\ud835\udf03\u2032\u2212 \ud835\udc57\ud835\udf14\ud835\udf0e( \ud835\udf070\ud835\udf0f 4\ud835\udf0b ) \ud835\udc4e\u222b 0 2\ud835\udf0b\u222b 0 { \ud835\udc5f sin (\ud835\udf03 \u2212 \ud835\udf03\u2032) \u2223\u2212\u2192\ud835\udc5f \u2212\u2212\u2192 \ud835\udc5f\u2032 \u22233 1 \ud835\udc5f\u2032 \u2202\ud835\udc48\ud835\udc67 \u2202\ud835\udf03\u2032 } \ud835\udc5f\u2032\ud835\udc51\ud835\udc5f\u2032\ud835\udc51\ud835\udf03\u2032 = \ud835\udc57\ud835\udf14\ud835\udf0e\ud835\udc35\ud835\udc671 (14) The closed form solution of (14) can\u2019t be obtained as it is basically an integro-differential equation. Hence, to solve (14), FDM can be applied by discretization of the disc into several elemental curvilinear zones by drawing \u2019n\u2019 equispaced concentric circles and \u2019m\u2019 equispaced radial lines as shown in Fig.5. Let, \u0394\ud835\udc5f be the elemental radial separation between any two consecutive circles (\ud835\udc5f\ud835\udc56 = \ud835\udc56\u0394\ud835\udc5f) and \u0394\ud835\udf03 be the elemental separation between any two consecutive radial lines (\ud835\udf03\ud835\udc57 = \ud835\udc57\u0394\ud835\udf03). Now, the integro-differential equation can be discretized around the point (i, j) by using central difference method, then the equation (14) becomes, (1 + 1 2\ud835\udc56 )(\ud835\udc48 \ud835\udc57 \ud835\udc56+1)\u2212 2\ud835\udc48 \ud835\udc57 \ud835\udc56 \u2212 (1\u2212 1 2\ud835\udc56 )(\ud835\udc48 \ud835\udc57 \ud835\udc56\u22121) + \ud835\udc48 \ud835\udc57+1 \ud835\udc56 (\ud835\udc56\u0394\ud835\udf03)2 \u2212 2\ud835\udc48 \ud835\udc57 \ud835\udc56 (\ud835\udc56\u0394\ud835\udf03)2 + \ud835\udc48 \ud835\udc57\u22121 \ud835\udc56 (\ud835\udc56\u0394\ud835\udf03)2 = \ud835\udc57\ud835\udf14\ud835\udf0e(\u0394\ud835\udc5f)2(\ud835\udc35\ud835\udc57 \ud835\udc671\ud835\udc56 + \ud835\udc35\ud835\udc57 \ud835\udc672\ud835\udc56) (15) Suppose, \ud835\udc48 \ud835\udc57 \ud835\udc56 = \ud835\udc48\ud835\udc58 where, k = (j-1)m + i \u21d2 (1 + 1 2\ud835\udc56 )(\ud835\udc48\ud835\udc58+1)\u2212 2\ud835\udc48\ud835\udc58 \u2212 (1\u2212 1 2\ud835\udc56 )(\ud835\udc48\ud835\udc58\u22121) + \ud835\udc48\ud835\udc58+\ud835\udc5a (\ud835\udc56\u0394\ud835\udf03)2 \u2212 2\ud835\udc48\ud835\udc58 (\ud835\udc56\u0394\ud835\udf03)2 + \ud835\udc48\ud835\udc58\u2212\ud835\udc5a (\ud835\udc56\u0394\ud835\udf03)2 = \ud835\udc57\ud835\udf14\ud835\udf0e(\u0394\ud835\udc5f)2(\ud835\udc35\ud835\udc671\ud835\udc58 + \ud835\udc35\ud835\udc672\ud835\udc58) (16) Consider the point P(i, j) to be the field point and \ud835\udc43 \u2032(\ud835\udc56\u2032, \ud835\udc57\u2032) to be the source point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003016_j.compstruct.2018.11.095-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003016_j.compstruct.2018.11.095-Figure1-1.png", + "caption": "Fig. 1. Diagram of the multiple tape design on a double-curved fuselage surface. (a) Current tape trajectory design in parallel planes, (b) New tape trajectory design using a geometry-induced application mechanism.", + "texts": [ + " One challenge is the lower conductivity of CFRP compared to metal [1]. To increase the conductivity, a thin metallic layer is placed on the exterior surface of the fuselage. For this purpose, preimpregnated tapes of lightning strike protection (LSP) are applied to the fuselage. Multiple single tapes (maximum width \u22480.8m) are aligned in order to cover the entire fuselage. The most challenging part of the tape design and application is the double-curved surface where the fuselage becomes more narrow ((1) see Fig. 1). The current approach for designing the aligned tapes (2) is to subdivide the fuselage in planar Sections (3) (see Fig. 1(a)). Here, it is necessary to variably steer the LSP material in the double-curved section. Known automation processes (fibre shearing and steering) are not adequate, despite the reduction of the tape width (width \u22480.6m). We introduce a new approach in which we neglect the constraint of a fixed tape orientation, as it is not necessary for LSP. Therefore, tapes (4) can be positioned more freely on the surface (see Fig. 1(b)). In addition, a focus on constraints that basically enable automation processes for wide tapes is given. The approach consists of two main issues, which are implemented in computer-based algorithms: \u2022 Geometry-induced application mechanism for one tape (Theory in Section 3, Experimental Investigation in Section 4). \u2022 Arrangement of multiple tapes for entire surface covering (Theory in Section 5, Calculation and Optimisation in Section 6). The developed application of LSP is a process with a specific set of constraints (wide tapes, ensured overlapping but no specific orientation)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003819_b978-0-12-812939-5.00002-1-Figure2.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003819_b978-0-12-812939-5.00002-1-Figure2.4-1.png", + "caption": "FIG. 2.4 Inertial reference frameOandnoninertial reference frameO0 rotatingwith instantaneous angular velocity\u03c9 ! .", + "texts": [ + " Well, first we need to realize that noninertial frame attached to joint 2 is both accelerating in translational sense as well as changing orientation of its axes with respect to axes of lab inertial frame. It is easy to deal with translational accelerations; one just needs to make sure that appropriate three-dimensional inertial force proportional to three-dimensional accelerations of joint 2 is considered in Eq. (2.57). But how to deal with the rotating axis? 2.3.6 Dynamics in Noninertial Accelerating and Rotating Coordinate Frame Consider inertial reference frameOand noninertial reference frameO0 depicted in Fig. 2.4. Quite generally, the origin of O0 can be accelerating and O0 axes can be rotating with instantaneous angular velocity \u03c9 ! . The radius vector of any point can be expressed in both frames and related to each other as r !\u00bc r ! oo0 + r !0 (2.106) The effect of angular velocity can be understood in the context of time evolution of radius vector in the lab frame. To simplify the problem, imagine that both r ! oo0 and r !0 are constant vectors. Furthermore, imagine that r ! oo0 \u00bc 0. Then for infinitesimally small period of time 4t, r " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.14-1.png", + "caption": "FIGURE 16.14 (a) Schematic of the keyhole cavity shape. (b) Typical top view of the weld pool during keyhole welding. (From Klemens, P. G., 1976, Journal of Applied Physics, Vol. 47, pp. 2165\u20132174. Reprinted by permission of American Institute of Physics, Melville, NY.)", + "texts": [ + " The width of the molten region determines the width of the weld seam. The simplest approximation to the keyhole configuration is that of a vertical cylinder (Fig. 16.13a). However, since the vapor pressure generally varies with depth, the radius is also a function of depth. The variation of pressure with depth results in outward flow of vapor, causing vapor to emerge from the end of the keyhole as an axial jet. Generally, the axis of the keyhole is bent in the direction of workpiece travel (Fig. 16.13b). One typical shape of the keyhole is shown in Fig. 16.14a, and has a constriction near the cavity outlet. Other shapes can also develop, depending on the welding conditions. Some are relatively short with walls that are convex outward and rounded at the bottom, while others are long and thin, concave outward and have sharp points at the bottom when blind. The shape shown in Fig. 16.14a tends to occur at relatively high welding speeds and larger penetration depths (or beam powers). Due to the motion of the heat source, the top surface of the molten metal develops a teardrop shape (Fig. 16.14b), with a typical length-to-width ratio of about ten, and the width tails off toward the back. At low welding speeds, the molten pool is almost circular in cross section, but the teardrop shape becomes more pronounced as the speed increases (see also Fig. 10.6). The width of the molten pool decreases at higher welding speeds, forcing the molten metal to move at a higher velocity through the narrow channel between the vapor and the solid boundary. When the power level is low enough, the vapor pressure produced in the keyhole is much smaller than the fluid dynamic forces of the molten metal surrounding the keyhole", + " Approximate Analysis An estimate of heat flow in keyhole welding can be done using the line heat source analysis discussed in Section 10.2 to obtain the temperature distribution, cooling rates, and peak temperatures, without taking into account the fluid flow (liquid, vapor, and plasma). The discussions in Section 10.3 and the preceding section provide a basis for analyzing the molten pool that surrounds the keyhole, but do not address the gaseous plume (vapor/plasma) within the keyhole. We add to these analyses by first looking at flow conditions in a plane parallel to the workpiece surface (Fig. 16.14b) and make the following assumptions: 1. All variations in the z-direction (normal to the workpiece surface) over a range comparable to the keyhole radius, rk, are small. This assumption may not be particularly valid close to the surface. 2. Welding speeds are high enough that adiabatic heating conditions exist ahead of the weld pool. In other words, heat conducted outward in the forward direction is not lost, but available to subsequently melt the material, unless it is conducted sideways, out of the path of the molten metal, faster than the advance speed. Using the nomenclature in Fig. 16.14b, the average speed at which heat is conducted sideways, ucs, may be approximated as ucs = \u03bas rk + wl (16.15) where rk = keyhole radius (mm), wl = average sideways width of the molten pool (mm), and \u03bas = thermal diffusivity of the solid (mm2/s). Now if the welding speed, ux, is greater than \u03bas rk , only a small fraction of the heat conducted out ahead of the molten pool will be lost. Thus, ux > \u03bas rk defines the \u201cadiabatic\u201d heating condition in laser welding. Weld Pool Length Ahead of the Keyhole, lf The length of the molten material between the keyhole and unmelted solid in front of it can be obtained by a simple heat balance that gives (cvolTm + Lmv)ux = kl Tv \u2212 Tm lf ( 1 \u2212 uxlf \u03bal ) (16", + "16) is based on the concept that the heat conducted through the molten metal provides enough thermal energy to advance the melt front at a speed ux. Since kl = cvol\u03bal, equation (16.16) reduces to the form lf = \u03bal ux ( Tv \u2212 Tm Tv + Lmv/cvol ) (16.17) from which we find that the length of the molten material between the keyhole and unmelted solid is inversely proportional to the welding speed. Average Side Width of the Molten Pool, wl If we divide the material around the cavity into four quadrants (Fig. 16.14b) and consider the first two quadrants, then in each of those two quadrants, sufficient heat must be conducted into the molten region to melt cold material at such a rate that the melt boundary in the quadrant moves with a speed ux. The average width of this melt is wl, and the average temperature, 1/2(Tm + Tv). The heat conducted across the liquid\u2013vapor interface, of projected area rk (assuming unit depth), under the influence of a temperature gradient Tv \u2212 Tm/wl, provides the energy required to heat and melt the new material", + " The resulting heat balance, neglecting heat conduction from the cavity into the molten region, as well as that conducted from the liquid into the solid, which are assumed to approximately cancel out, gives rk[\u03b2vLvol + cvol(Tv \u2212 Tm)] = Lmv(D \u2212 rk \u2212 wl) (16.26) where D is the maximum width of the molten zone (mm), Lvol is the latent heat of vaporization (or condensation), per unit volume of liquid (J/m3). The length of the molten pool can be approximated by considering the distance l that the beam advances during the time tD that it takes for the melt to attain the width D, that is, the time it takes for heat to be conducted from the rear of the cavity over a distance D. l (see Fig. 16.14b) is then given by l = uxtD = ux D2 \u03bal (16.27) Thus, from equations (16.26) and (16.27), we have l = ux rk 2 \u03bal ( 1 + \u03b2vLvol + cvol(Tv \u2212 Tm) Lmv + wl rk )2 (16.28) Pressure Variation in the Keyhole For the purposes of this discussion, pressure within the keyhole is considered to vary only along the length of the keyhole, that is, in the vertical direction and is uniform in a horizontal section. As indicated earlier, the cavity is kept open by an excess pressure of vapor in the keyhole, Pv, (the plasma plume) and ablation pressure, Pab" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001649_012093-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001649_012093-Figure4-1.png", + "caption": "Figure 4. Result of motion synthesis to the end point of the synthesized path.", + "texts": [ + " Input of array data \u03a81, \u03a82, and \u03a83 specifying the knowledge base of the past experience; 2 is the calculation of via points AP1 and AP2 specifying the position of a synthesized path for the OL centre point movement, these via points providing the absence of the path and forbidden regions P1 and P2 intersection (provision of minimum offset to forbidden regions); 3 is the determination of parameters specifying icons for projection of arm working envelope 0 i based on the known coordinate points B\u04211, B\u04212 [11]; 4 is the recognition of points AS, AP1, AP2 and AE belonging to the working envelop [11]; 5 is the calculation of vector q\u03941 and values xS, yS specifying an optimal rest position of the android arm and body (at Us = max), relative to the manipulation object based on the parameter zi and array \u03a81; 6 is the analysis of value compliance q\u03941, point coordinates AS, AP1, AP2, AE, B\u04211, B\u04212 with array values \u03a82, which prescribes deadlock; 7 is the motion synthesis to change the configuration type and determine a new value q [9]; 8 is the calculation of the generalized velocity vector ),,,( 21 iMMMM qqqQ based on the range of motion minimization criterion on the path sections prescribed by the segments ASAP1, AP1AP2 and AP2AE [2,10]; 9 is the determination of the arm configuration condition assigned by the values and forbidden regions (here the assumption is made that \u0394qi \u2248 iMq ), iMq are the components of the vector QM; 10 is the construction of the next configuration or iNii qqq ; 11 is the determination of maximum values ki max based on qi and the array \u03a83 [12]; 12 is the change of values ki = ki + 1 used in the vector equation (3); 13 are the values ki satisfying the maximum values ki max; 14 is the synthesis of motions with invariant OL centre position [9]. The motion synthesis using weighting factors of generalized velocity values. The arm motion to change the type of configuration; 15 is the calculation of the generalized velocity vector ),,,( 21 iNNNN qqqQ (3) [8]; 16 is the target point reached in the prescribed segment; 17 is the change in the value of the next configuration nk= nk + 1; 18 is the result output of the motion synthesis along the complete path section determined by the segments ASAP1, AP1AP2 and AP2AE. In figure 4 the results of the motion synthesis using the knowledge bases to transfer the manipulation object from the point \u0410S to the point \u0410E are shown. In figure 4 the motion synthesis with the type of configuration being changed and using weighting factors of generalized velocities values is shown on the path section prescribed by the points \u0410S\u0410P1. 6 1234567890 \u2018\u2019\u201c\u201d The developed algorithm for virtual control of android motion using the developed knowledge bases allows the complex evaluation of current situations to be performed and from this the optimal logic choice to be implemented. Minimal total change of generalized coordinates occurs when making this choice" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003738_978-3-030-04975-1_7-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003738_978-3-030-04975-1_7-Figure4-1.png", + "caption": "Fig. 4. Defining section ui", + "texts": [ + " The determination of the resulting clearance in mesh Di is necessary to determine the actual distance ui and the value of intertooth Fi. The manufacturing tolerances of the individual elements influencing the creation of intertooth clearance Di are shown in Fig. 3. Therefore, the following manufacturing tolerances can be distinguished: \u2022 tolerance for the gear outline - Tzh; \u2022 tolerance for the roller - Tr; \u2022 tolerance for the rollers arrangement radius - TRg; \u2022 tolerance for the angular arrangement of the rollers - TuR; \u2022 tolerance for the eccentric - Te. Based on Fig. 4, the following can be specified: \u2013 distance ui ui \u00bc rw2 yozri cos# \u00fewi \u00f07\u00de \u2013 distance wi wi \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi xzh xozri\u00f0 \u00de2 \u00fe yozri yzh\u00f0 \u00de2 q \u00f08\u00de The real outline of the hypocycloidal gearing, with deviation dzh from the theoretical outline can be described in relation to the X02Y system as follows: xzh \u00bc q z 1\u00f0 \u00de cos g\u00fe k q cos z 1\u00f0 \u00deg\u00fe g\u00fe dzh\u00f0 \u00de cos g k cos z 1\u00f0 \u00degffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 k cos zg\u00fe k2 p \u00f09\u00de yzh \u00bc q z 1\u00f0 \u00de sin g k q sin z 1\u00f0 \u00deg\u00fe g\u00fe dzh\u00f0 \u00de sin g\u00fe k cos z 1\u00f0 \u00degffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 2 k cos zg\u00fe k2 p where: dzh \u2013 deviation of the theoretical hypocycloidal outline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000779_icit.2009.4939600-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000779_icit.2009.4939600-Figure5-1.png", + "caption": "Fig. 5. Contour plot of no load flux density of flameproof motor.", + "texts": [ + " The resulting differential equations are: [ ] { } { } [ ] [ ] { } [ ] { } { } \u03a9=\u00d7\u2207\u00d7\u2212\u2207+ \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 \u2202 \u2202+\u22c5\u2207\u2207\u2212\u00d7\u2207\u00d7\u2207 inAVV t AAvAv e 0\u03c3\u03c3 \u03c3 (2) [ ] [ ] { } [ ] { } { } \u03a9 =\u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u00d7\u2207\u00d7+\u2207\u2212 \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 \u2202 \u2202\u22c5\u2207 in AVV t A 0\u03c3\u03c3\u03c3 (3) [ ] ( ) ( ) ( ){ }3,32,21,1 3 1 3 1 vvvvtrve ++== (4) where V is electric scale potential, \u03c3 is conductivity, \u03bd is reluctivity, \u03a9 is conducting region. Since in the end region, the motor has 3-D magnetic flux. It is necessary to use 3-D finite element analysis (FEA) for accurate determination of eddy current field. The model for field solution in the end region is shown in Fig. 4. It has two parts, circuit part and finite element part. The circuit part is used as magnetic field excitation and the finite element part is used for field analysis. Fig. 5 shows the no load magnetic distribution in the end region. C. 3D Solid Source Conductor Inductance Calculation The typical structure of the solid source conductor is illustrated in Fig. 6. The magnetic flux density B is a vector. It can be expressed as in (5) kzyxRjzyxQizyxPzyxB ),,(),,(),,(),,( ++= (5) where P, Q and R have consecutive first order partial derivative. The magnetic flux \u03c6 can be expressed by surface integral as in (6) \u222b\u222b \u2211 ++= RdxdyQdzdxPdydz\u03c6 \u222b\u222b \u2211 ++= dSRQP )coscoscos( \u03b3\u03b2\u03b1 dSBdSnB n\u222b\u222b\u222b\u222b \u2211\u2211 =\u22c5= (6) where \u2211 is a directed surface in the vector field, n is normal unit vector, Bn is the magnetic flux density B projection on the normal vector of the surface\u2211 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000021_imece2009-11165-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000021_imece2009-11165-Figure3-1.png", + "caption": "FIG. 3: ARC HEAT SOURCE ON SURFACE OF ROTATING CYLINDER", + "texts": [ + "org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2009 by ASME \u03be\u03be \u03c0\u03c1 \u03be dZK cv qT LX LXp 22 0e +=\u0394 \u222b + \u2212 \u2212 (9) with the Bessel function of the second kind and zero order, K0. The variable \u03be is used for integration. The dimensionless coordinates X, L, Z are defined by Eq. 8. Ling et al. [12] have transferred the equation of heat conduction Eq. 5 onto circular systems. An arc heat source (included angle \u03b1\u03d5\u03b1 +\u2264\u2264\u2212 , heat flux per unit area q(\u03d5)) affects the surface of a rotating cylinder, see Fig. 3. The radius of the cylinder is a. A heat loss by convection occurs on the cylinder\u2019s surface outside the area affected by the heat source. The equation of heat conduction yields \u239f\u239f \u23a0 \u239e \u239c\u239c \u239d \u239b \u2202 \u0394\u2202 = \u2202 \u0394\u2202 + \u2202 \u0394\u2202 + \u2202 \u0394\u2202 \u03d5 \u03c9 \u03ba\u03d5 TT rr T rr T 111 2 2 22 2 (10) For the described cylindrical system the boundary conditions are ( ) ( )\u23a9 \u23a8 \u23a7 >\u2265 \u2264\u2264\u2212 \u2212 = \u2202 \u2202 \u03b1\u03d5\u03c0 \u03b1\u03d5\u03b1 \u03d5 \u03d5 ,ahT q r Tk (11) with the heat transfer coefficient h. For the surface temperature at ar = a solution of Eq. 10 in normalized form with the shown boundary conditions is ( ) ( ) ( )[ ] ( )\u222b \u2212 + = \u03b1 \u03b1 \u03c8\u03c8\u03d5\u03c0\u03c8\u03c8 \u03d5 dKuf u ,2 (12) with ( ) ( ) ( ) ( ) ( )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002860_s11668-018-0561-y-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002860_s11668-018-0561-y-Figure3-1.png", + "caption": "Fig. 3 (a, b) Close-up views of corresponding fracture surfaces on a broken axle, showing three distinctly identifiable regions: (I) initial quenching crack, (II) fatigued region, and (III) final rupture", + "texts": [ + " The broken components were examined visually as well as metallographically, while at least three broken axles were sectioned longitudinally for examining the profile of induction-hardened case. As stated above, the fracture pattern of almost all the broken axles was identical and is shown in Fig. 2. M. R. Shad (&) F. ul Hasan Department of Mechanical Engineering, University of Central Punjab, Johar Town, Lahore 54000, Pakistan e-mail: dr.rizwan@ucp.edu.pk A close-up view of the fracture surface shown in Fig. 3 shows that three regions on the fractured surface were distinctly identifiable: 1. Initial crack, which (as discussed later) was essentially within the induction-hardened case. 2. Progressive fatigued region, showing typical \u2018beach marks.\u2019 3. Final rupture which appears to be essentially shear. The depth of the \u2018origin\u2019 of the crack in almost all the broken axles was observed to be about 3\u20134 mm. Some of the broken axles were longitudinally sectioned and macroetched to examine the profile of induction-hardened case. Figure 4 shows the profile of hardened case, which was found to be in compliance with the specification as shown in Fig. 1. Figure 4 also shows that the original crack had formed in the \u2018neck\u2019 region of the axle and that the depth of the initial crack (which is clearly visible in Fig. 3) was essentially the same as the depth of the induction-hardened case. It was also noticeable that this initial crack was normal to the surface at the location where it was formed, a feature which is typical of quench cracking [1]. The \u2018curved\u2019 path of the fatigue crack may also be noted. Microstructure shown in Fig. 5, which was taken from a location indicated by arrowhead, shows that the crack was normal to the surface and was intergranular, i.e., features which are consistent with quenching cracks [1, 2]", + " However, the most interesting feature of the present fracture was that the fatigued portion of the fractured surface was \u2018curved,\u2019 i.e., the fatigue crack had gradually changed its orientation as it progressed (a rare phenomenon). Clearly, the initial quench crack although had provided an origin to the fracture, its orientation was not subjected to the maximum tensile stress. Hence, the crack, as it progressed, gradually acquired an orientation where the plane of the crack was normal to the operative tensile stress [1, 3]. The final rupture (indicated by (III) in Fig. 3) on the basis of its orientation in most of the broken axles has been regarded as shear. The proportion of this area when considered in relation to the total fractured surface again indicates a low nominal stress on the axles during service [3]. The overall analysis clearly suggests that the formation of quench cracks during induction hardening was the root cause of the failure of drive axles which were otherwise subjected to fairly low nominal stress during service. The attention thus had to be focused on why the quench cracking during induction hardening had occurred in some of axles which were induction-hardened during the winter period of November to January/February" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000087_12.819551-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000087_12.819551-Figure2-1.png", + "caption": "Fig 2 shows the pinion head-cutter. Head-cutter surface of pinion is represented as:", + "texts": [], + "surrounding_texts": [ + "Fig1 shows the cone of gear head-cutter. The derivation of the gear tooth surface is based on the following procedure: Step1: the cone surface is represented as follow: g g 2 g e g g g g 2 g g 2 ( sin )cos ( , ) ( cos )sin cos r u r u r u u \u03b1 \u03b8 \u03b8 \u03b1 \u03b8 \u03b1 \u23a1 \u23a4\u2212 \u23a2 \u23a5= \u2212\u23a2 \u23a5 \u23a2 \u23a5\u2212\u23a3 \u23a6 ur (1) Here: gu and g\u03b8 are the surface coordinates; 2\u03b1 is the blade angle; gr is the radius of gear head-cutter. Step2: by the equation of meshing and the coordinate transformation, the gear tooth surface is represented as follow: 2 2 g g( , )r r \u03b8 \u03c6= r r (2) Here: g\u03c6 is the rotation angle of cradle in the process of gear generation." + ] + }, + { + "image_filename": "designv11_92_0001986_s0025654418010065-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001986_s0025654418010065-Figure4-1.png", + "caption": "Fig. 4.", + "texts": [ + " Thick line denotes the graph of K(\u03b1), the fine lines represent the graphs of the function tan(\u03d5 \u2212 \u03b1) for certain values of \u03d5. Let \u03d5 be somewhat bigger than \u03c0/2. Obviously, both isolated solutions will be displaced somewhat, and the value of \u03b1 increases. A set of uninsulated solutions disappears, isolated branch \u03b13(\u03d5) separates from its left edge, root \u03b12(\u03d5) moves to the right, approaching the value of \u03c0 for \u03d5 \u2192 \u03c0. With increasing \u03d5 roots \u03b11(\u03d5) and \u03b13(\u03d5) approach each other and, at some value of \u03d5 = \u03d5\u2217 merge into one \u03b1\u2217, and then vanish, which is reflected in the right-hand side of Fig. 4 a. The left part of the picture is completed symmetrically in accordance with property b). This \u201cprocessing\u201d of Fig. 3 and the use of formulas (3.3) allow us to construct the dependences u\u0304(\u03d5) and v\u0304(\u03d5) on steady-state rotation regimes. These dependencies are qualitatively represented in Figs. 4 b and 4 c. Points A, B, B1, C, C1, which will be useful in constructing phase portraits in paragraph 5, are marked on them. MECHANICS OF SOLIDS Vol. 53 No. 1 2018 3.2. Stability of Steady Regimes Entering, as usual, deviations from the steady-state values of the variables x = u \u2212 u\u0304, y = v \u2212 v\u0304, z = \u03b8 \u2212 \u03b8\u0304 = \u2212(\u03b1 \u2212 \u03b1\u0304) we can write down the first-approximation equations by virtue of equations (2", + " 7 b shows a phase portrait in the case where gravity is present. In the sectors noted above, the straightness of the phase trajectories is preserved. But the lower sector is divided by the curve AB, consisting of uninsulated fixed points. In addition, on the line v = u tan \u03b11 there is another isolated fixed point C (stable node). Obviously, in region u < 0 there exist a fixed point C1 (u\u0304C1 = \u2212u\u0304C) and an arc AB1. The curve points of the BAB1 correspond to the non-isolated modes identified in Fig. 4 b and 4c by vertical segments. Stationary point A corresponds to a vertical steady-state descent mode (without rotation) with velocity v\u0304A = \u22121/ \u221a CD(\u03c0/2). The velocities in steady descending regimes corresponding to points B and C are equal: u\u0304B = w\u0304(\u03b1\u0303) cos \u03b1\u0303, v\u0304B = \u2212w\u0304(\u03b1\u0303) sin \u03b1\u0303, u\u0304C = w\u0304(\u03b11) cos \u03b11, v\u0304C = \u2212w\u0304(\u03b11) sin \u03b11, w\u0304(\u03b1\u0303)2 = 1 CD(\u03b1\u0303) sin \u03b1\u0303 + CL(\u03b1\u0303) cos \u03b1\u0303 = sin \u03b1\u0303 CD(\u03b1\u0303) , w\u0304(\u03b11)2 = 1 CD(\u03b11) sin \u03b11 + CL(\u03b11) cos \u03b11 . It follows from Fig. 4 c and 7 b that v\u0304A < v\u0304B < v\u0304C < 0. This demonstrates the effectiveness of using a rotochute in comparison with a parachute. 4.3. Characteristic Features of the Phase Portrait for a Rotational Body with Inclined Blades For \u03d5 \u2208 [0, \u03c0 \u2212 \u03d5\u2217) \u222a (\u03d5\u2217, \u03c0] the solution of equation (3.4) is unique \u03b1\u0304 \u2208 [0, \u03b10] \u222a [\u03c0 \u2212 \u03b10, \u03c0], and the steady-state motion regime is unique. On the phase plane it corresponds to a singular point u = u\u0304, v = v\u0304 \u2013 the only attractor to which all other phase trajectories contract" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001113_978-1-84800-239-5_62-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001113_978-1-84800-239-5_62-Figure3-1.png", + "caption": "Figure 3. Free body diagram of the gears forming a planet gear set", + "texts": [ + " The efficiency of the system can be obtained from the following expressions: 1 1 2 6 1 1 2 6( ) /( )s out H out out H outT T T T (12) where the output component of differential branch is the carrier, so Tout=TH1, 1 1out HT T . The output component of differential branch is the ring, so 2 6outT T , 2 6outT T . After getting the external torque values applying on each component, the gear mesh and bearing forces of each deck can be calculated using the static equilibrium. The expressions of the gear mesh and bearing forces are as following according to the relationship between torque and force using the diagram shown in the Fig. 3. s r H sp pr pH s r p H T T TF F F r r n r where Fsp is the gear tangent mesh force between the sun and planet gear. Fpr is the force between the ring and planet gear. FpH is the planet bearing force applied on the carrier by the planet. Here, rs and rr are the pitch radius of s and r defined from the center of the gear to the pitch point as shown in the Fig. 3. rH is defined from the rotational center to the center of a planet. Applying these theories on the power flow planet gears, the gear tangent mesh and planet bearing forces can be get according to the following the expressions: ( 1, 4), ( 3,6) , ( 1, 2)j j Hi spi pri pHi j j pi Hi T T TF j F j F i r r n r (13) where Fspi, Fpri and FpHi represent the gear tangent mesh force of sun and planet gear, planet and ring, the planet bearing force in the ith branch respectively. rj is the pitch radius of each gear and rHi is defined from the rotational center to the center of a planet in the ith branch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000939_sice.2008.4655035-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000939_sice.2008.4655035-Figure3-1.png", + "caption": "Fig. 3 A coordinate system of the 2D unicycle.", + "texts": [ + " PR0001/08/0000-2229 \u00a5400 \u00a9 2008 SICE - 2230 - MMS612E-500, MITSUBISHI PRECISION Co., Ltd Therefore, we derive a motion equation of a system which considered a force and a torque to the system by the operator using the Projection method. We develop a dynamic unicycle simulator, using the equation of motion, which prepare an environment for analysis, training, teaching, and assist of the unicycle operation. 2. MODELING OF A UNICYCLE We perform a modeling of a unicycle, including the operator\u2019s input, in Fig.3[6, 7]. Note that we consider a pedal mass of the unicycle for considering the operator\u2019s input. 2.1 Modeling of the 2D unicycle A generalized coordinate xu2D and a generalized ve- locity vu2D are defined to be xu2D = [\u03b8w \u03b8s \u03b8pL \u03b8pR xw zw xs zs xpL zpL xpR zpR]T vu2D = [ \u02d9\u03b8w \u03b8\u0307s \u03b8\u0307pL \u03b8\u0307pR x\u0307w z\u0307w x\u0307s z\u0307s x\u0307pL z\u0307pL x\u0307pR z\u0307pR]T . A tangent velocity is q\u0307u2D = [ \u02d9\u03b8w \u03b8\u0307s \u03b8\u0307pL \u03b8\u0307pR]T . A generalized mass matrix Mu2D be defined to be Mu2D = diag(Iw, Is, Ip, Ip, mw, mw, ms, ms, mp, mp, mp, mp). Where Iw, Is, Ip are a principal moment of inertia of a wheel, a saddle and paddles of the unicycle, and mw,ms,mp are a mass of the wheel, the saddle and the peddales. Let sin\u03b1, cos \u03b1, tan\u03b1 denote S\u03b1,C\u03b1,T\u03b1. An input value to the peddles FL, FR and saddles \u03c4s given by sensors in fig.1 and fig.3. Thus a generalized is force as: hu2D = [\u2212Cws \u201c \u03b8\u0307w \u2212 \u03b8\u0307s \u201d , \u03c4s + Cws \u201c \u03b8\u0307w \u2212 \u03b8\u0307s \u201d , \u03c4pL, \u03c4pR, 0,\u2212gmw, 0,\u2212gms, \u2212 S\u03b8pLFL,\u2212C\u03b8pLFL \u2212 gmp, \u2212 S\u03b8pRFR,\u2212C\u03b8pRFR \u2212 gmp]T , where Cws is a viscous friction between the wheel and the saddle. From a relation of a gravity point of the wheel, the saddle and peddles, we get 8>< >: xw = rw\u03b8w , zw = rw xpL = lcS\u03b8w + xw , zpL = lcC\u03b8w + zw xpR = \u2212lcS\u03b8w + xw , zpR = \u2212lcC\u03b8w + zw xs = rsS\u03b8s + xw , zs = rsC\u03b8s + zw . From these equations, we get a constraint matrix Cu2D for which Cu2D vu2D = 0, Cu2D = h Cu2D1 Cu2D2 i Cu2D1 = 2 6666666664 rw 0 0 0 0 0 0 0 lc cos \u03b8w 0 0 0 \u2212lc sin \u03b8w 0 0 0 \u2212lc cos \u03b8w 0 0 0 lc sin \u03b8w 0 0 0 0 rs cos \u03b8s 0 0 0 \u2212rs sin \u03b8s 0 0 3 7777777775 Cu2D2 = 2 6666666664 \u22121 0 0 0 0 0 0 0 0 \u22121 0 0 0 0 0 0 1 0 0 0 \u22121 0 0 0 0 1 0 0 0 \u22121 0 0 1 0 0 0 0 0 \u22121 0 0 1 0 0 0 0 0 \u22121 1 0 \u22121 0 0 0 0 0 0 1 0 \u22121 0 0 0 0 3 7777777775 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000701_ichve.2008.4773917-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000701_ichve.2008.4773917-Figure3-1.png", + "caption": "Figure 3. Infinite length line charges and its mirror", + "texts": [ + "00 \u00a92008 IEEE Separate the surface of every splitting conductor into 24 arc segments with same length, and number all nodes from 0 to 23, as shown in Fig.2. Set the surface charges density of every node be \u03c3i, i=0, 1, 2,\u2026, then the surface charges density of any point on the splitting conductor surface which have a angle \u03b8 is: (1) )1( 24 2 24 2 if 24 2 24 )1(2 )( 1 ii\u03c0ii ii +<< \u2212 + \u2212+ = + \u03c0\u03b8\u03c3\u03c0\u03b8\u03c3 \u03b8\u03c0 \u03b8\u03c3 Thus, the surface charges on splitting conductor can be regarded as many infinite length line charges with a line density r\u03c3(\u03b8)d\u03b8, as shown in Fig. 3. According to electrostatic field theory, the potential of any point (x, y) produced by infinite length surface charges with an arc segment width and its mirror to the ground is: (2) ln)( 2 1),( 11 )1( 24 2 24 2 1 2 0 ++ + +== \u222b iiii i i ccd R R ryxd \u03c3\u03c3\u03b8\u03b8\u03c3 \u03c0\u03b5 \u03d5 \u03c0 \u03c0 (3) ))sin(())cos(( 2 0 2 01 yyxrxR \u2212++\u2212+= \u03b8\u03b8 (4) ))sin(())cos(( 2 0 2 02 yyxrxR +++\u2212+= \u03b8\u03b8 Here: (x0, y0) is the coordinates of splitting conductor circle centre, r is radius of splitting conductor, ci and ci+1 are coefficient that can be calculated by actual coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000124_pime_auto_1964_179_013_02-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000124_pime_auto_1964_179_013_02-Figure1-1.png", + "caption": "Fig. 1. Beam axle geometry", + "texts": [ + " It is thus necessary to apply to the system a sufficient and necessary number of co-ordinates to describe all possible movements, to describe the energy of the system in terms of these co-ordinates and hence to substitute the appropriate values in the equation. One suspension, the beam axle, will be discussed in sufficient detail to demonstrate the approach. Results will be quoted for the other systems. A complete derivation of Lagrange\u2019s equation will be found in reference (I)*. Reference (2) discusses the application of the method to engineering problems and gives examples. The body is assumed to be connected to the axle by two springs (Kv) which are separated as shown, Fig. 1. The lateral flexibility of the springs, bushes and supports is combined into a single lateral stiffness (K,) effect represented by the horizontal spring between body and axle. The body is considered to rotate about 0,, \u2018the instantaneous centre of roll\u2019 of the body. In the case of the beam axle * References are given in Appendix II. Proc Instn Mech Engrs 1964-65 Vol179 PI 2A No 3 at UNIV OF CINCINNATI on June 4, 2016pad.sagepub.comDownloaded from 100 J. R. ELLIS this is assumed to be at the height at which a line joining the front and rear spring attachment points to the body crosses a vertical line through the centre of the road wheel when the vehicle is viewed in side elevation", + " f 2 f ) + Q x +2C cos +[R,d-R&-(k-zi7) cos 41 = k,(h,+&)+Qw mi-2KU(Z-x)+ 2 k ~ - 2C sin2 +(Z-i) mw-Kl( Y-w)+2k,(aRo+w) ie-22Ku1,2(O- a) +2kIRo(aRo+w)+2kuL2~ +2C[RZ2&- RlR2d+(Y-th)R2 cos $1 = h ( h i +X2)Ro+kuW1f-f,f)+ Qa For small movements about a position of equilibrium the products of variables may be neglected and cos 8 --f 1, sin 8 + 8. The linearized equations are expressed in matrix form, Table 1. The left-hand side of the matrix provides complete information about the natural frequencies of vibration and the stability of the suspension. It is usual to expand this term and obtain a polynomial the roots of which will give the natural frequencies of the system and which can be tested for stability. In the case of vehicle suspensions, however, the examination of stability is not considered to be necessary since inspection of Fig. 1 will reveal that the most probable case of instability would be if the moment produced by the c.g. of the body about the roll centre, 0, during roll (MgH sin 6) was sufficient to overcome the resistance of the main springs in roll. No practical vehicle is likely to be unstable because of this feature. Proc Instn Mech Engrs 196465 When the equations are written in matrix form it is convenient to use the operation form D( ) = d( )/dt. D2( ) = d2( )/dt2. A steady state position under the action of external forces is reached when the derivatives of each and every variable become zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003470_amcon.2018.8614963-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003470_amcon.2018.8614963-Figure3-1.png", + "caption": "Fig. 3 Correlation between the cutting tool and the tool axis vector", + "texts": [ + "2 Machine tools application service module requirements data analysis & visualization Platform operation & maintenance Device connectivity & data management Inform ation security operations support Data Analysis Automated testing Paid service Data visualizationComplex Event Engine App and service management Data management & storage Equipment management SCM (supply chain management) CRM (Customer relationship management) ERP (Enterprise resource planning) MES ( Manufacturing Execution System) Mechanical End Controller Layer Acceleration gauge, temperature, current PLC Industrial Computer Switching board CNC Controller Production management digitization (from industry 2.0 to 3.0) Build National IoT PaaS (PaaS) CNC Machine Tool Factory Machine Electronic process equipment Peripherals AI functions (quality control, process optimization, predictive diagnosis, accuracy optimization) Fig.3 Smart manufacturing solutions [2] Conclusions Develop an overall industrial solution to provide a full range of intelligent application services is very important. Through software and hardware integration, we will increase the added value of products, develop smart manufacturing's overall solutions, and minimize price competition scenarios. Intelligent equipment that has undergone trials and verifications in the market promotes stand-alone, complete line or whole plant output. References [1] Ministry of Economic Affairs(MOEA), Promotional program of smart machinery industry, May, 2016", + " The data represented in Table 1 does not contain the tool axis orientation which should be determined further. Generally, tools are divided into ball mill, end mill and bull nose mill. This paper would use the ball mill as an example to explain the mathematical module. The five-axis tool path file provided by Mastercam will include the tool tip vector 1 T x y zQ Q Q Q , the contact point vector T x y z 0N N N N , and the contact point vector 1 T x y zP P P P . By multiplying the known ball radius R with the contact point vector as shown in Fig. 3, the tool center point coordinate can be calculated. Plus, with the use of known contact point coordinate and the vector relationship between the known contact point coordinate and the contact point to the tool center, the tool center point vector 1 T x y zC C C C can be found. Since the tool was a ball mill, the theoretical tool axis vector was the unit vector between the tool tip and the tool center point. Therefore, the tool vector K could be calculated from the tool tip coordinate and the tool center point coordinate, based on the equation below: R R P N QK (1) Once the tool tip position and tool orientation are known, the NC program including three linear movements (X, Y, Z) and two rotary movements (B, C) can be obtained by homogenous coordinate transformation and inverse kinematics [4]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000028_icems.2009.5382869-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000028_icems.2009.5382869-Figure1-1.png", + "caption": "Fig. 1. Static and dynamic eccentricities.", + "texts": [ + " Therefore, the circulating current is conceivably generated in the parallel winding. However, while many reports have been presented about unbalanced magnetic force due to rotor eccentricity, but there have been very few concerning the circulating current of the parallel winding. This paper details the main factors why the rotor eccentricity degrades the PMSM performance degradation by using the results from a magnetic field analysis and measurements taken from the parallel armature stator windings. Figure 1 shows two kinds of the rotor eccentricity: (a) static and (b) dynamic classes. For the static rotor eccentricity, the stator axis is parallel to the rotor axis. For the dynamic rotor eccentricity, the stator axis is not parallel to the rotor axis. In this report, we described the influence of the static eccentricity on the PMSM characteristics. Figure 2 shows an example of a PMSM machine: (a) a structure with an off-center rotor arrangement and (b) parallel Y connections of the armature stator windings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003481_cistem.2018.8613602-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003481_cistem.2018.8613602-Figure2-1.png", + "caption": "Fig. 2. Component selection", + "texts": [ + " Radio control: a radio remote control (RC) is an instrument for controlling a vehicle remotely. It is usually equipped with two potentiometers for controlling the power of the engine and steering. It is equipped mainly by a transmitter and receiver equipped with a high frequency module. Radio communication requires a minimum of four channels associated to Roll, Pitch, yaw and altitude. The RC used for our model is a turnigy tgy-i10 which contains 10 channels with a transmitter and receiver of 2.4 GHz frequency. The selected components are represented by fig 2 and more explained in [6]. Mass: the mass is very important in a drone, if it is light, it gains reliability and flight time. Using an electronic scale we can identify the total weight of our drone and also the weight of each component. Motor dynamics: Brushless DC motor modeling can be approximated to that of a DC motor and thus its transfer function is of second order approximated to a first order. One of the methods to identify the motor dynamics is to apply a voltage setpoint and measure, with a magnetic speed sensor, the allure given by the motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001719_cyberi.2018.8337561-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001719_cyberi.2018.8337561-Figure7-1.png", + "caption": "Fig. 7. Virtual balancing model in MD Adams", + "texts": [ + " SOFTWARE MODEL OF BAM This section presents software model of the real balancing arm and describes in details the steps of its creation and verification. Built Balancing Arm Model was also implemented using a computer simulation where we used the SolidWorks environment for the original 3D layout. The creation of the computer model followed the practical realization. This means creating a computer model according to the dimensions of the actual device. We have kept all the necessary dimensions to achieve the most accurate computer duplicate, Fig. 7, compared to the real model. Both models have the same component dimensions as well as the same material thicknesses. Building the model was completed step by step, by creating the individual parts that we have connected together like during the real model construction. By checking the dimensions, we have made sure that we have a technical copy of the original with which we can work in MD Adams. In this program environment, we can simulate the behavior of the real system. The first step after importing the 3D model from SolidWorks, is to set and check the relationships between the individual components of the software model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002497_978-3-319-99620-2_12-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002497_978-3-319-99620-2_12-Figure7-1.png", + "caption": "Fig. 7. Equivalent strain distribution on pressure vessel with optimized geometry", + "texts": [ + " Idea was to allow higher stresses in pressure vessel to evaluate plastic strain, but still significantly lower than tensile strength; according to the materials database [30] tensile strength is between 500 and 750 MPa. After RSO had been done, table of optimization gave three \u201ccandidates\u201d (i.e. combinations of parameters) based on specified goal and constraint. Trade off study was then conducted to reveal feasible points and, eventually, the best candidate was selected. Original model was updated with new values of parameters and numerical simulation was carried out once again. Values of strain in the area of interest on the optimized pressure vessel are given in Fig. 7. Results obtained in experiment and FEA showed good agreement, i.e. both approaches indicated the same area of the highest strain in the vicinity of the one of the nozzles. This was the sign that all elements of the numerical model (geometry, load, boundary conditions) were well defined. Response Surface Optimization (RSO) method was used to optimize geometry of the pressure vessel parts (shell, head and nozzles). Response surfaces (RS) are functions of different nature where the output parameters are described in terms of the input parameters", + " Objective of the optimization was minimization of the vessel\u2019s mass by varying the thicknesses of nozzles\u2019 walls as well as of vessel head and cylindrical surface. Using Response Surface Optimization method mass of the vessel was reduced from 3.642 to 3.383 kg, which gives the total reduction of 1.036 kg (considering that one fourth of the vessel was modelled). Thicknesses of nozzles\u2019 walls were increased by 0.85 mm (nozzle 1) and 3.2 mm (nozzle 2), while the thickness of the cylindrical surface was reduced by 0.3 mm (from 1.5 to 1.2 mm). Thickness of the vessel head remained almost the same: initial was 2.2 mm while final somewhat higher \u2212 2.23 mm. Figure 7 shows that strain value around nozzle 1 on optimized vessel is almost doubled (0.388% compared to 0.20% on original model and measured in experiment) which is expected considering the fact that thickness of cylindrical surface is now reduced, and the higher value of stress was allowed in optimization (new value of maximum equivalent stress is 315 MPa compared to 200 MPa on original vessel). Nevertheless, the main goal is achieved, and further investigations will show are these values of stress and strain acceptable (from the point of view of vessel integrity) and if they are, is there more room for further mass reduction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002221_978-3-319-99270-9_27-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002221_978-3-319-99270-9_27-Figure2-1.png", + "caption": "Fig. 2. Ball bearing model", + "texts": [ + " The model for the backup bearing, duplex pair of angular contact ball bearing, is developed on the basis of the ball bearing model introduced by Sopanen and Mikkola [13]. The bearing forces are calculated based on the bearing geometry, material property and the deformation of the bearing in the dropdown. In the model, the relative displacement between races is given by: erj \u00bc ex cos bj \u00fe ey sin bj etj \u00bc ez wx sin bj \u00fewy cos bj Rin \u00fe rin\u00f0 \u00de \u00f02\u00de where ex;y;z represent the relative displacements of the bearing races (Fig. 2), and wx and wy show the tilting of the inner race in x, y-axis. bj is the attitude angle of j th ball. Thus, the distance between the races can be calculated as follows: Dj \u00bc rout \u00fe rin Rin \u00fe rin \u00fe erj Rout \u00fe rout cosuj \u00f03\u00de where uj is the contact angle. uj \u00bc tan 1 etj Rin \u00fe rin \u00fe erj Rout \u00fe rout ! \u00f04\u00de where Rin and rin represent the inner race radius and inner race groove radius, correspondingly. The following equation expresses the total elastic deformation of the inner ring. dtotj \u00bc dj Dj \u00f05\u00de The Hertzian contact theory has been used for the calculation of the contact force in the dropdown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000549_phealth.2009.5754830-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000549_phealth.2009.5754830-Figure5-1.png", + "caption": "Fig. 5 Miniaturized prototypes realized in LTCC illustrating the integrated spiral coil antennae (arrow 1) and nanoporous membrane (arrow 2), with the pressure transducer mounted on the reverse side.", + "texts": [ + " a spiral coil on the surface of the 3 x 7 mm implant carrier) has been connected to the ASIC with power supplied from a coil at 7 mm distance to verify sufficient efficacy of the wireless interface. The back telemetry with the real scale antenna has been tested with a load modulation implemented with discrete components. A later ASIC implementation of the load modulation has been successfully tried with a macro scale coil antenna. The sensors and electronics developed for the prototype will be integrated into a miniaturized device suitable for subcutaneous implantation. The miniaturized device (fig. 5) exhibits comparable sensor architecture to the prototype. Emerging membrane technologies providing narrower distribution around the MWCO such as track etched polymer, anodic alumina or thin film nanoporous silicon glass membranes will be investigated. The nanoporous membrane measures here 1 x 1 mm2 in extent and is attached to a 2x2 mm2 large silicon frame. The membrane encloses a reference chamber measuring 1.2x1.2x0.5 mm3 which is more than 260 times smaller than the prototype. The differential pressure transducer (of similar size geometry as the membrane/silicon frame) is located on the opposing side" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000408_s1068798x09100220-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000408_s1068798x09100220-Figure3-1.png", + "caption": "Fig. 3. Verification of the model in Eq. (3) according to the Box\u2013Cox criterion: (1) midpoint of the interval 0.09.", + "texts": [ + " Three remainders fall beyond the upper limit of the confidence interval. According to the user\u2019s wishes, the program indicates these points in the plan matrix with repetitions: 41 \u2013 A = C = D = +1.00 and B = E = \u20131.00; 142 \u2013 C = +1.81 and A = B = D = E = 0.00; 151 \u2013 A = B = C = D = 0.00 and E = \u20131.81. Analysis of the model of the least-square estimates according to the Box\u2013Cox criterion confirms that in fact the operator \u03bb = 1 for Eq. (3) is not within the confidence interval; its best value must be regarded as \u03bb = +0.09 (Fig. 3). However, taking account of the convenience of the subsequent transformation of the model y\u03bb relative to the function y, which is ultimately of interest to the user, the pro- 1 5, ), 1 6, ) R\u0302a1; t\u0302 p 1( ), 10 50, ; R\u0302a2; R\u0302z2; t\u0302 p 2( ), 20 40, . t\u030230 1( ) 8.845 1.654AD 2.29CE\u2013 3.027DE\u2013+= \u2013 1.57A2 0.641A3, %.+ models of the least-squares estimates (a) and the most-probable estimates (b). t\u030230 1( ) RUSSIAN ENGINEERING RESEARCH Vol. 29 No. 10 2009 INFORMATION BASE FOR REGULATING THE GRINDING 1061 gram recommends the adoption of \u03bb = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000779_icit.2009.4939600-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000779_icit.2009.4939600-Figure9-1.png", + "caption": "Fig. 9. Distribution of core loss of a flameproof motor at no load.", + "texts": [ + " All coefficients displayed in (14) to (16) can be deduced from the data measured by the rotational core loss tested by curve fitting (13) and (14). The core loss with elliptical B is predicted from the alternating and circularly rotating core losses by: ( ) altBrotBfe PRPRP 21\u2212+= (17) where majB BBR /min= is the axis ratio, minB and majB are the major and minor axes of the elliptical B locus. In the laminated steel motor, the flux density vector rotates in an elliptical pattern, but within the lamination plane. The distribution of stator core losses at no load is shown in Fig. 9. It can be seen that core loss in the yoke is greater than that in the teeth because the amplitude changing of magnetic field in the yoke is greater. The objective of this paper is to present a method to analyze the distribution of eddy current in the end region of flameproof motor. This has been successfully accomplished. Though this method needs to calculate the inductance of each stator coils, it has less degree of freedom and computing time (compared to magnetic nodal method using directly coupling method)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000905_speedham.2008.4581079-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000905_speedham.2008.4581079-Figure1-1.png", + "caption": "Fig. 1. Phasors diagram for the synchronous motor", + "texts": [ + " The internal angle,\u03b8 , is usually chosen as parameter of load; this angle is defined as being the angle between the phasor of the terminal voltage and the phasor of the e. m. f. induced by the inductor field. III. EQUIVALENT CIRCUIT HAVING THE INTERNAL ANGLE \u03b8 AS PARAMETER The following voltages equation written in complex for one phase of the armature winding results by considering that the motor is supplied by a voltage having the root-mean-square U and the frequency f IRIjXILjILjUU qaqdade ++++\u2212= \u03c3\u03c9\u03c90 , (1) with f\u03c0\u03c9 2= (2) and the other quantities have known meaning [1, 3, 4]. In accordance with this, the phasors diagram depicted in the figure 1 can be plotted and the following relations are obtained by means of the projections on the d, q axes: qedd RIUIfLU ++= 02cos \u03c0\u03b8 RIIfLU qq \u2212= \u03c0\u03b8 2sin (3.a,b) The influence of the armature reaction phenomenon and the magnetic core losses are neglected further on. By noting with: 0,, eNN UfU the rated supplying voltage, the rated frequency and the e. m. f. induced by the inductor field at rated speed and with Nu UU /=\u03b1 (4) the signal factor for voltage, Nf ff /=\u03b1 (5) the signal factor for frequency and Nee UUk /0= (6) the excitation factor, the equations system (3) is written in the form: qddfNefNu RIIXUkU +=\u2212 \u03b1\u03b1\u03b8\u03b1 cos ; 513 978-1-4244-1664-6/08/$25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002890_j.matpr.2018.08.148-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002890_j.matpr.2018.08.148-Figure2-1.png", + "caption": "Fig. 2. Testing rig.", + "texts": [ + " The spring manufacturers are obligated to perform this type of tests before introduction the product to the market, so the springs are designed to meet the regulations requirements. However, the real operation conditions of the vehicles equipped in this type of the suspension system are more demanding than the test assumptions. An approach to the mounting method of the springs described in the standard is also quite flexible. This also influence the fatigue strength. In order to determine the number of cycles, which the analyzed spring is able to resist, a specific test stand was built. It enabled to conduct a fatigue test of a complete object in a controlled way (fig. 2). 26762 Sta\u0144co M., Iluk A., Dzia\u0142ak P./ Materials Today: Proceedings 5 (2018) 26760\u201326765 The type of the mounting of the spring strongly influence the stiffness and the operation of the element. The authors managed to recreate the mounting of the leaf spring on the vehicle. In the middle part of the object the deflection was restricted by the clamp. The endings of the element were placed on the rollers. It enabled unlimited length change during the spring deflection. The examination was conducted continuously, by the control of the displacement (bend deflection of the spring f=145mm) with frequency equal 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001831_0142331217739688-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001831_0142331217739688-Figure6-1.png", + "caption": "Figure 6. Bearing test platform 1.", + "texts": [ + " Compute the features of the test samples by the tech- nology described in step 2 to step 4. 7. Recognize the classes of the test samples using the trained SVM. In this section, two real bearing data sets are utilized to validate the effectiveness of our proposed method. Bearing data set 1: The bearing data set is obtained from the Western Reserve University Bearing Data Center website. This data set, sampled from a motor test platform, has become a standard data set used to verify the efficiency of a new fault diagnosis method. As shown in Figure 6, the test platform consists of a motor (left), a torque transducer/encoder (centre), a dynamometer (right), and control electronics (not shown). It can generate four types of data including normal data, inner race fault data, outer race fault data, and ball fault data. The dimensionality of each sample is 1024, and the size of each kind of data is 100. The one-dimensional signals in the time domain are shown in Figure 7. In this experiment, we mainly analyse the performance of MLSLLE in terms of feature extraction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001861_978-3-319-79111-1_42-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001861_978-3-319-79111-1_42-Figure3-1.png", + "caption": "Fig. 3 Evaluation of fitness value", + "texts": [ + " For a feasible architecture of the mechanism, the values l1 and l2 have to be lower than LM and the angles \u03b11 and \u03b21 to respect the sizes of the U and S joints. Also, it has been considered that the sum of radii r and R to be lower than the sum of l1 and l2 in order to keep the lateral dimensions of the mechanism as low as possible. The objective function evaluates two components. The first is represented by the area of the 2-dimension illustration of the angles of the sun presented in the Fig. 2 that are not covered by the workspace of the mechanism. Referring to the Fig. 3 the uncovered areas S1 and S2 from the angles of the sun represent one part of fitness value, the sum of S1 and S2 being a surface (a scalar value). The second component of the fitness value is the total area of the FP and MP (a scalar value). By subtracting the second term from the first one it is assured that the PM covers as many sun angles as possible, but keeping its geometrical dimensions to minimum as seen in the Eq. (5). fit= S1 + S2 \u2212 \u03c0\u00f0R2 + r2\u00de \u00f05\u00de The GA has been implemented in the Global Optimization Toolbox from Matlab [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002481_gt2018-76823-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002481_gt2018-76823-Figure1-1.png", + "caption": "FIGURE 1. VIEW OF THE TEST RIG", + "texts": [ + " The rotating test rig exploited in this experimental campaign is placed at the Laboratory for Technology for High Temperature (THT) at the Department of Industrial Engineering of the University of Florence. A soundproof cabin contains the test facility that is completely remotely controlled by means of control pulpits placed outside, assuring the safety of the operators. A ventilation system equipped with filters allows for air recirculation avoiding overheating of the instrumentation. As clearly shown in Figure 1, the rotating frame is placed over a steel plate which is decoupled from the base by means of elastic joints in order to eventually dump vibrations. An electric spindle provides the rotation to the system, it can reach a maximum velocity of 15000 RPM and a maximum torque of 30 Nm. A high precision bearing-less torque-meter connects the motor to the driving shaft, that is supported by a pair of angular contact ball bearings mounted in \u201cO\u201d arrangement (load lines diverge along the bearing axis) contained in a sealed bearings casing equipped with an oil jet cooling system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000302_pi-b-2.1962.0194-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000302_pi-b-2.1962.0194-Figure4-1.png", + "caption": "Fig. 4.\u2014Diagram showing the element of force PQ in S at / and the corresponding P'Q' in S' at /'.", + "texts": [ + " 1 and 2 showing a 2-dimensional movement, the position of a point O' was plotted with different sets of co-ordinates according to whether its position was seen by an S' observer to be fixed, or by an S observer to be moving. In the 4-dimensional problem now to be treated, it is helpful to picture x, y, z for a fixed system S, and x', y', z' for a moving system S', not separately, but as a single set of co-ordinates with common scales. In this way, and in contrast with Figs. 1 or 2, every instantaneous position must have different plots depending on whether it is seen by an S or an S' observer. Each plot must also be labelled with the time to which it relates. Fig. 4 is drawn in this manner. It shows the instantaneous position of an element PQ of a line of electric force, with its inherent velocity vector c. P and Q are points seen in S at /, and P' and Q' are the same points but seen in S' at t'. The system S' moves with velocity jSc in the x-direction with respect toS. The Figure has been simplified by making P and P' simultaneously coincide, so that they both lie at the origin of the common x, y, z and x', y', z' axes and /' is synchronized with t at that origin" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000817_iciea.2009.5138297-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000817_iciea.2009.5138297-Figure2-1.png", + "caption": "Figure 2. Six sectors of three-phase input voltage and the combinations of DC bus voltage", + "texts": [ + " The Strategy of IMC Supposing the expression of input voltage is: cos( ) 2cos( ) 3 2cos( ) 3 a m i b m i c m i u V t u V u V t \u03c9 \u03c0\u03c9 \u03c0\u03c9 = = \u2212 = + (9) We can divide three phase input voltage into six parts so that there is a positive DC voltage in DC side and we can get a biggest output voltage and a smaller switch loss. Each sector has the same performance that is absolute value of one phase is largest and others have opposite polarity. The period of each PWM can be divided two section and two positive polarity voltage are output in the two sections. So two effective space vector voltage of rectifier can be created in a period and zero vectors can\u2019t be produced just like Fig.2. For example, in the third sector ub is positive and maximal, then phase b is always conducted in rectifier circuit. Phase a and phase c are modulated. So the mean value of DC voltage in one period is: dc cb cb ab abU d u d u= \u22c5 + \u22c5 (10) 1cb abd d+ = (11) In the equation, ,bc bad d is duty ratio of cbu and abu in one period. 1). In Figure.1 b), when phase c is modulated, Scn and Sbp will conduct. Output voltage in the rectifier is Udc=ub- uc. The duty ratio of phase c and phase b is: cos / coscb c bd \u03b8 \u03b8= \u2212 (12) 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001736_j.ifacol.2018.04.023-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001736_j.ifacol.2018.04.023-Figure1-1.png", + "caption": "Fig. 1. An asteroid and a tethered particle.", + "texts": [ + " (1982, 1993); Levin (2005); Rodnikov (2006), etc. Dynamics of tethered systems attached to an asteroid are studied intensively by Misra (2012, 2014, 2016); French et al. (2014); Misra (2016a). For the asteroid and the massive point of comparable masses, the conditions of existence and stability for steady motions are studied by Burov (2016). Numerical investigations of two tether-connected masses rise, probably, to Ivanov et al. (1995). Consider a tethered system (TS) that consists of asteroid A and a spacecraft Q of mass m, see Fig. 1. The spacecraft is connected to the point P of the asteroid surface by a weightless flexible inextensible tether of length : |PQ| \u2264 . We assume that the asteroid moves as an Euler top and that the motion of the spacecraft Q does not affect the asteroid dynamics. Let OX\u03b1X\u03b2X\u03b3 be an absolute reference frame (ARF) with the origin in the asteroid center of mass O and let Ox1x2x3 be the mobile reference frame (MRF), formed by the central principal axes of inertia of asteroid A. The direction cosines of ARF axes with respect to MRF are \u03b1 = ( \u03b11 \u03b12 \u03b13 ) , \u03b2 = ( \u03b21 \u03b22 \u03b23 ) , \u03b3 = ( \u03b31 \u03b32 \u03b33 ) ", + " (1982, 1993); Levin (2005); Rodnikov (2006), etc. Dynamics of tethered systems attached to an asteroid are studied intensively by Misra (2012, 2014, 2016); French e al. (2014); Misra (2016a). For the asteroid and the massive point of comparable masses, the con itions of xistence and stability for steady otions are studied by Burov (2016). Numerical investigations of two tether-connected masses rise, probably, o Ivanov et al. (1995). Consider a tethered system (TS) that consists of asteroid A and a spacecraft Q of mass m, see Fig. 1. The sp cecraft is connected to the p int P of th asteroid surface by a weightl ss flexible inextensible te r of length : |PQ| \u2264 . We assume that the asteroid mov s as a Euler top and that the motion of the spacecraft Q does not affect the asteroid dynamics. Let OX\u03b1X\u03b2X\u03b3 be an absolute reference frame (ARF) wi h the origin in the aster id cent r of mass O and let Ox1x2x3 be the mobile ref ence frame (MRF), formed by the central principal axes of i ertia of asteroid A. The direction cosines of ARF axes with respect to MRF are \u03b1 = ( \u03b11 2 3 ) , \u03b2 = ( \u03b21 2 3 ) , \u03b3 = ( \u03b31 2 3 ) ", + " (1982, 1993); Levin (2005); Rodnikov (2006), etc. Dynamics of tethered systems attached to an asteroid are studied intensively by Misra (2012, 2014, 2016); French et al. (2014); Misra (2016a). For the asteroid and the massive point of comparable masses, the conditions of existence and stability for steady motions are studied by Burov (2016). Numerical investigations of two tether-connected masses rise, probably, to Ivanov et al. (1995). Consider a tethered system (TS) that consists of asteroid A and a spacecraft Q of mass m, see Fig. 1. The spacecraft is connected to the point P of the asteroid surface by a weightless flexible inextensible tether of length : |PQ| \u2264 . We assume that the asteroid moves as an Euler top and that the motion of the spacecraft Q does not affect the asteroid dynamics. Let OX\u03b1X\u03b2X\u03b3 be an absolute reference frame (ARF) with the origin in the asteroid center of mass O and let Ox1x2x3 be the mobile reference frame (MRF), formed by the central principal axes of inertia of asteroid A. The direction cosines of ARF axes with respect to MRF are \u03b1 = ( \u03b11 \u03b12 \u03b13 ) , \u03b2 = ( \u03b21 \u03b22 \u03b23 ) , \u03b3 = ( \u03b31 \u03b32 \u03b33 ) ", + " (1982, 1993); Levin (2005); Rodnikov (2006), etc. Dynamics of tethered systems attached to an asteroid are studied intensively by Misra (2012, 2014, 2016); French et al. (2014); Misra (2016a). For the asteroi and the massive point of comparable asses, the conditions of existence and stability for steady motions are studied by Burov (2016). Numerical investigations of two tether-connected masses rise, probably, to Ivanov et al. (1995). Consider a tethered system (TS) that consists of asteroid A and a spacecraft Q of mass m, see Fig. 1. The spacecraft is connected to the point P of t e asteroid surface by a weightless flexible inextensible tether of le gth : |PQ| \u2264 . We assume that the asteroid moves as an Euler top and that the motion of the spacecraft Q does not affect the asteroid dynamics. Let OX\u03b1X\u03b2X\u03b3 be an absolute reference frame (ARF) with the origin in the asteroid center of mass O and let Ox1x2x3 be the mobile refere ce frame (MRF), formed by the central principal axes of inertia of asteroid A. The direction cosines of ARF axes with respect to MRF are \u03b1 = ( 1 2 \u03b13 ) , \u03b2 = ( 1 2 \u03b23 ) , \u03b3 = ( \u03b31 \u03b32 \u03b33 ) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure11-1.png", + "caption": "Fig. 11 Schematic of the coordinate system", + "texts": [ + " U1637206 and 51705311), the SAST Project (Grant No. SAST2017-079) and the State Key Laboratory of Mechanical System and Vibration of Shanghai Jiao Tong University (Grant No. MSVZD201709). Declaration of Conflicting Interests The authors declare no potential conflicts of interest with respect to the research, authorship, and/or publication of this article. Nomenclature \u03b1 Axial deflection parameter \u03b2 Radial deflection parameter E Modulus of elasticity Appendix A circular thin plate is studied in the coordinate system, as shown in Fig. 11, in which r is the radial direction and w is the axial direction. According to related knowledge on the axisymmetric bending of circular thin plates in Ref. [28], the maximum axial displacements under various conditions are obtained, as shown in Table 4. Although bearing stiffness is variable, parameters \u03b1 and \u03b2 are calculated based on Table 4 in this study. As shown in Table 4, the two parameters are almost invariable. Therefore, we do not need to perform finite element calculation all the time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure2-1.png", + "caption": "Fig. 2. Two phases of mechanisms joint", + "texts": [ + " This metamorphic joint is combined by link a\u223c d and three revolute joints with axis 1, axis 2 and axis 3. Link a is fixed in base or moving platform, link d is connected to other links. In the metamorphic joint, rotational axis 1, rotational axis 2 and rotational axis 3 are perpendicular to each other in the initial state. The output degree of freedom is two when revolute joint of axis 1 or axis 2 is locked. There have two phases when changing position of link b and link c through locking revolute joint of axis 1 or axis 2 in Fig. 2. Two typical phases of metamorphic joints are demonstrated with the link a fixed with either base or platformas shown in Fig. 2. Two phases are confirmed with respect to the relative position between the link b and link c. When revolute joint of axis 2 is locked and link b is coincided with link c and perpendicular to link a, as show in Fig. 2(a), it gives e1 phase. In this phase, rotational axis 1, rotational axis 2 and rotational axis 3 are perpendicular each other. Link d has two rotations along rotational axis u1 and w1 directions in local coordinate system o1\u2212u1v1w1 in this phase. When revolute joint of axis 1 is locked and link b is fitted perpendicularly to link c and link a as in Fig. 2(b), it gives e2 phase. In this phase, rotational axis 1 and rotational axis 3 are parallel and perpendicular to rotational axis 2. Then link d has two rotations along axis u1 and v1 directions. Two phases are presented pursuant rotating the link to suitable position, which including two rotations in every phase via locking the rotational joint. Metamorphic joint take on differential phases with changing the position of link, then there exist two phases could be chosen, which phases from one to another, that satisfy the characteristic of metamorphic, then joint is metamorphic joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000881_pi-b-2.1962.0229-Figure15-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000881_pi-b-2.1962.0229-Figure15-1.png", + "caption": "Fig. 15.\u2014Bent and folded notches.", + "texts": [ + " The gap between the top of the fin and the rudder mass-balance acted as a parasitic notch, resonant within the required frequency band, and prevented the desired current flow around the periphery of the tip of the tail structure, so that 436 BURBERRY: PROGRESS IN AIRCRAFT AERIALS there was no rearward radiation as the radiation patterns of Fig. 14 indicate. This effect has been used to give directivity to a notch aerial in an application described in a later Section of the paper. Where structural limitations prevent the use of a straight notch, it is possible to obtain the required length by bending or folding the notch, as Fig. 15 indicates. In the limiting case, one surface of the notch can be reduced to a thin diaphragm between the two legs, but there is some loss of bandwidth with this form. Johnson2 has shown that the resistance at the open end of the notch is of the order of 2 000 Q, and the parameters of the notch may be determined by treating it as a transforming section. The portion between the open end and the feed point is considered as a parallel-plate transmission line of varying characteristic impedance depending on the cross-section and separation of the conducting surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003825_1754337119831107-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003825_1754337119831107-Figure6-1.png", + "caption": "Figure 6. Arrangement for measuring tare drag on the axial sting. (a) Air flow nozzle (17 mm \u00d8). (b) Pellet. (c) Pellet axial sting (1 mm \u00d8; not attached to pellet). (d) Pellet transverse sting (1 mm \u00d8). (e) Pellet elevation adjusting mechanism. (f) Pellet two-dimensional translation system (mechanism not shown). (g) Pellet rotation mechanisms. (h) Axial sting attached to load cell for measuring tare drag force on axial sting (see Figure 5 for attachment details).", + "texts": [ + " The load cell was mounted on a three-dimensional (3D) positioning stage, which allowed the pellet axis to be aligned accurately with the nozzle axis, with the aid of an aligning jig. When the pellet was mounted, the air flow imposed drag forces, not just on the pellet but also on the pellet support sting. Therefore, the force measured by the load cell, was the gross drag value, which included the air drag force on the pellet itself plus the \u2018tare\u2019 drag due to the air jet impinging on the sting. In order to evaluate the net drag on the pellet, the tare drag on the sting was measured (after each measurement of the gross drag), using the additional arrangement shown in Figure 6. An identical pellet was mounted, using a transverse sting (item D in Figure 6), to a support system (items E, F, G) which was independent of the load cell. An axial \u2018dummy\u2019 sting (item C), not connected to the pellet, was placed in position and attached to the load cell of Figure 5. Thus, the arrangement of Figure 6 allowed the tare drag on the dummy axial sting to be measured and subtracted from the overall (gross) drag, resulting in the required value of net drag force acting on just the pellet. This pellet overall net drag force was then used to calculate the pellet drag coefficient as shown in equation (5) Cd = drag force 1 2 rU2A \u00f05\u00de where drag force is the net drag force and A is the pellet cross-sectional area at the pellet head rim. Measurement of pressure distribution around the pellets The pressure distribution on the surface of the four pellets numbered 2 to 5 in Figure 2 was measured as illustrated in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002873_smasis2018-7915-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002873_smasis2018-7915-Figure4-1.png", + "caption": "FIGURE 4: COORDINATE SYSTEM AND DISCRETIZATION OF A SINGLE SEGMENT OF AAFB WITH ELASTIC SUPPORTS", + "texts": [ + " The activation of the patch actuators causes a mechanical strain in the circumferential direction of each pad by the inverse piezoelectric effect, which changes the radius and the center of curvature of each shell. Further detailed properties of the functional model are described in [4]. The following sections present a model based on the Ritz method for the curved plates to simulate the shape control of the adaptive segments in the functional model. The model is then validated using experimental results for single supporting shells (without elastic supports) and the functional model. Fig.4 shows a schematic representation of a single support shell and the associated piezoelectric patch actuator (consisting of two insulation layers and the piezoelectric ceramic in between) with elastic supports for further analysis. The elasticity can be related linearly to the electric field, with subsequent definition of inverse and direct piezoelectric effects in piezoelectric materials: { Dp } = [ep] { \u03b5p } +[\u03b7p] { Ep } (5) { \u03c3p } = [Cp] { \u03b5p } \u2212 [ep] T {Ep } (6) 5 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings", + " The shells have two planes of symmetry, namely the \u03be1 \u2212 \u03be3 plane and \u03be2 \u2212 \u03be3 plane. As a result, the following conditions must be applied in the basis functions so that the existing symmetry planes are correctly considered: u\u03be1 (\u03be1,\u03be2) = u\u03be1 (\u03be1,\u2212\u03be2) u\u03be1 (\u03be1,\u03be2) =\u2212u\u03be1 (\u2212\u03be1,\u03be2) u\u03be2 (\u03be1,\u03be2) =\u2212u\u03be2 (\u03be1,\u2212\u03be2) u\u03be2 (\u03be1,\u03be2) = u\u03be2 (\u2212\u03be1,\u03be2) u\u03be3 (\u03be1,\u03be2) = u\u03be3 (\u03be1,\u2212\u03be2) u\u03be3 (\u03be1,\u03be2) = u\u03be3 (\u2212\u03be1,\u03be2) \u03c8\u03be1 (\u03be1,\u03be2) = \u03c8\u03be1 (\u03be1,\u2212\u03be2) \u03c8\u03be1 (\u03be1,\u03be2) =\u2212\u03c8\u03be1 (\u2212\u03be1,\u03be2) \u03c8\u03be2 (\u03be1,\u03be2) =\u2212\u03c8\u03be2 (\u03be1,\u2212\u03be2) \u03c8\u03be2 (\u03be1,\u03be2) = \u03c8\u03be2 (\u2212\u03be1,\u03be2) (14) Assuming a single bearing shell in different main parts, as shown in Fig.4, it can be seen that there are changes in stiffness at the borders of each area. In addition, actuation with the corresponding piezoelectric patches is only possible in the area II. Now the approximation functions for both areas are set up separately and linked with consistency conditions: uI \u03be1 (\u03be1,\u03be2) = uII \u03be1 (\u03be1,\u03be2) uI \u03be2 (\u03be1,\u03be2) = uII \u03be2 (\u03be1,\u03be2) uI \u03be3 (\u03be1,\u03be2) = uII \u03be3 (\u03be1,\u03be2) \u03c8 I \u03be2 (\u03be1,\u03be2) = \u03c8 II \u03be2 (\u03be1,\u03be2) (15) Due to the symmetry, it is sufficient to set up approximation functions for the areas I and II, because the deformation in area III is symmetrical to that in area I", + " In mechanical systems, this variation formulation directly follows the Hamilton principle, which can be expressed in the following form for static models: \u2202L \u2202q = 0 (17) where the Lagrangian L consists of the potential energy in the system and the work of external forces: L =W (q)\u2212\u03a0np,p (18) 7 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/15/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use The energy stored in the elastic suspension can be considered both as potential energy or as work of external forces, depending on where the system boundary is assumed and should be added to the terms of energy in piezoelectric and non-piezoelectric elements in the single supporting shell. Back to the schematic representation of a single support shell in the functional model (Fig.4), the solid joints react to forces in the circumferential direction of the bearing shells as well as to forces in the radial direction or to acting moments. Thus, the flexure joints can be simplified at two ends of each shell with three linear springs, namely in \u03be3 direction, perpendicular to \u03be3 and about the \u03be1 axis. The potential energy in elastic joints can be expressed as following: \u03a0J = \u222b {uJ} 0 {FJ} d {uJ}= 1 2 {uJ}T [CJ ]{uJ} (19) where {uJ} is the displacement vector at the end of a single shell (where the joints are connected) and [CJ ] is the stiffness matrix in flexure joints" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure29-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure29-1.png", + "caption": "Fig. 29 Control flow of wired humanoid", + "texts": [ + " Two cables actuating the knee joint are wrapped around the pulley, the axis of which coincides with that of knee joint. By virtue of this structure, the moment arm on knee joint torque is constant regardless of knee joint angle. As a consequence, the wide range of joint movement can be achieved. Each edge point of three cables actuating the ankle joints has the cylinder-shaped cable guide as shown in Fig. 28. The axis of each cylinder is parallel to the pitch axis. This wire guide prevents the cables bending below its minimum bend radius and diminishes the moment arm variation around the ankle joints. Figure 29 shows the control flow of the wire-driven biped robot. The blocks in Fig. 29a are the same as the controller described previously in Sect. 3; hence, the details of these blocks are omitted. The control procedures specific to the wiredriven robot are the blocks in Fig. 29b, which are designed to enhance the joint servo performance of wire drive system as in Fig. 30. This controller has two types of control mode such as \u201cflexible\u201d and \u201cprecise.\u201d The control mode is changed depending on the wire tensional force, which is selected with the rightmost switch in Fig. 30. For instance, in the supporting phase, the robot needs to stay on its feet and to generate high tension on a pulling cable of stance leg joints; hence, the \u201cprecise\u201d mode is selected. In contrast, the joints of swing leg are controlled in the \u201cflexible\u201d mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002073_ccdc.2018.8407818-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002073_ccdc.2018.8407818-Figure5-1.png", + "caption": "Fig. 5 Mathematical model of ATT occupying", + "texts": [ + " The discovery angle and distance of the incoming torpedo relative to the ATT also changing, the speed of the ATT changing as well. During the maneuver of the ATT, There is a moment, the required velocity direction of the favorable advance angle determined by the relative position of the incoming torpedo and the ATT, starboard angle and speed ratio is similar to that of the ATT at this time. Therefore, this time can be regarded as the starting stage of direct search for ATT and then go into the direct search phase. As shown in Figure 5,the ATT while turn the maneuver after navigate steady. Suppose that ATT is located at point E, the incoming torpedo is at point D, the included angle between the torpedo velocity direction and the oz axis is Q. The distance between point D and E is tgR , ATT speed is tV , the incoming torpedo speed is gV , The starboard angle of incoming torpedo relative to the ATT is mQ , the optimal shooting advance angle \u03d5 determined by relative position of the incoming torpedo and ATT can be obtained by equation (2)-(5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003529_0954406218812119-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003529_0954406218812119-Figure7-1.png", + "caption": "Figure 7. Parametric 3D model of the detachable beam.", + "texts": [ + " Optimization target parameters (OTP) is set as the target parameter of the particle swarm optimization algorithm, and changing the value of X and optimizing the value of OTP, the value of OTP to the maximum value is set. N is half of the total length of the beam, and L is the part of the beam that binds to the column. N and L are set as the value of optimization constraint functions in the particle swarm optimization algorithm. The values of N and L constrain the range of values of X together. Parametric 3D model of detachable beam is shown in Figure 7. The parametric three-dimensional model of the detachable beam is used to determine the optimization target by changing the optimization variables. Based on the particle swarm optimization algorithm, the optimal design of the detachable beam is carried out. X represents the position of the joint surface of the detachable beam. Because the position of each X is a potential solution, in the optimization process, any position where X is located requires a given joint surface parameter (including the stiffness and damping of the joint surface)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000422_amr.76-78.713-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000422_amr.76-78.713-Figure1-1.png", + "caption": "Fig. 1 Generalized coordinates for Reynolds", + "texts": [ + " It is the perquisite to calculate the friction resistance between the shear stress and contact bodies. So, the effects of shear stress of oil film in mill bearing has been theoretically analyzed in detail, which has great meaning to further understand and grasp the lubricating performance of contact components under heavy load, as well as to prolong the bearing service life. Calculating Model on Shear Stress of Lubricating Film. When oil-film bearing works, the lubricating oil will be brought into the wedge clearance to form into hydraulic film by the rotation of roll neck. Fig.1 shows the coordinate system of Reyonlds equation. During the whole calculation process, bearing was regarded as elastic body 1, roll as elastic body 2 and bearing chock as elastic body 3. Meanwhile, it is supposed that the rotating roll and still bearing had no axial misalignment. Mathematical Model on Shear Stress of Lubricating Film. Shear stress formula can be deduced from the definition of Newtonian fluid. ,x y u v z z \u03c4 \u03b7 \u03c4 \u03b7\u2202 \u2202 = = \u2202 \u2202 (1) According to the rotation relationship between the bearing and the roll [6], velocity boundary conditions can be expressed as below, where 0 , hU U are respectively the velocity components in x direction at 0z = and z h= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003909_ecace.2019.8679324-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003909_ecace.2019.8679324-Figure1-1.png", + "caption": "Figure 1: An illustration of pitch control system", + "texts": [ + " Moreover, the results will help the researchers to find the characteristics of these controller. The remaining of the paper is arranged as follows: section 2 represents system modelling, section 3 describes controller design, section 4 evaluates the performance of the proposed controller and section 5 concludes the paper. The aim of this section is to represent the generalized modeling of pitch angle regulation of aircraft. The system is modeled by linearized model of aircraft with considering different flight condition parameters. Fig. 1 presents the illustration of pitch control system. Here, Xd, Yd and Zd present the aerodynamic force. The parameters \u03c3, \u03d5 and \u03b3a present the pitch, roll and elevator deflection angle respectively. Fig. 2 presents the force, moment and velocity components. Here, L, D and E present the force, moments and velocity element respectively. The angular rates are shown by p, q and r. The parameters p, q and r represent the roll, pitch and yaw axis. The velocity of roll, pitch and yaw axis can be represented by u, v and w respectively. The system modelling data is used from: NAV IONa [1]. The parameters of dimensional derivative Q = 35.80lb/ft2, QSc\u0304 = 38495ft.lb and (c\u0304/2uf) = 0.015 are used for this paper. Table 1 shows the longitudinal directional stability parameters. Assuming the aircraft can maintain uniform altitude and velocity. Therefore, the thrust force and drag force eliminates each other. Besides, The weight and lift correspondence each other. Considering the pitch angle remains unchanged. From Fig. 1 and Fig. 2, the rigid body equation can be obtained. The lateral directional movement has been defined as X \u2212mgS\u03c3S\u03c3 = m(u\u0307 + qv \u2212 rv) (1) D = \u2212Iy q\u0307 + rq(Iy \u2212 Ix) + Ixz(p 2 \u2212 r2) (2) The statement (1), (2) and (3) are nonlinear. The disturbance theory has been used to linearize the equations. Solving the aircraft problem, it is required to consider the following assumption as: rolling rate p = \u03d5\u0307 \u2212 \u03a5\u0307S\u03c3 , yawing rate q = \u03c3\u0307C\u03d5+\u03a5\u0307C\u03c3S\u03d5, pitching rate r = \u03a5\u0307C\u03c3C\u03d5\u2212\u03c3\u0307S\u03d5 , pitch angle \u03c3\u0307 = qC\u03d5 \u2212 rS\u03d5, roll angle \u03d5\u0307 = p+ qS\u03d5T\u03c3 + rC\u03d5T\u03c3 and yaw angle \u03a5\u0307 = (qS\u03d5+rC\u03d5)sec\u03c3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.56-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.56-1.png", + "caption": "Fig. 11.56 Simulated deformation of a tire during severe cornering. Reproduced from Ref. [50] with the permission of SAE", + "texts": [ + "55 compares the prediction and measurements of the fore\u2013aft force of the tire, showing good agreement up to a large slip ratio. The vehicle/tire interaction needs to be investigated in designing a suitable tire for a vehicle, because the tires are the only points of contact between the vehicle and the road. However, FEA has not been conducted for the tire/vehicle behavior under extreme operating conditions because it is difficult to analyze the large deformation of rotating tires. Fukushima et al. [50] simulated the tire/vehicle behavior of the J-turn using LS-DYNA and compared the prediction with measurements. Figure 11.56 shows the finite element model of a vehicle and tires and the deformation of the right-front tire near the contact patch in a cornering event. Although the transient behavior cannot be observed in an experiment, a simulation can provide a transient contact pressure distribution of a tire in cornering. We thus gain insights from the prediction that we cannot get experimentally. The side force and self-aligning torque of a tire are nonlinear with respect to the slip angle and slip ratio. Furthermore, these nonlinearities vary with the road temperature, traveling distance, wear and load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001499_978-981-10-5768-7_49-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001499_978-981-10-5768-7_49-Figure2-1.png", + "caption": "Fig. 2 Experimental set-up. The test disk (1) is fixed in a cylindrical holder and lowered a distance (d) into the melt pool (3). The temperature of the disk is measured with a thermocouple (1) whose wires runs through a steel tube insulated with a ceramic material (2). The chamber above the disk (5) is depressurized with a vacuum pump connected to the top plate (4)", + "texts": [ + " Two sets of five disks were built in two different powder bed fusion (PBF) machines, as shown in Table 1, to simulate a casting mould. Each disk had six groups of venting slots with 8 slots in each group. The nominal slot widths of each venting slot group is given in Table 1. The diameter of the test disk is 80 mm and the thickness is 5 mm. The disks were built with the slots aligned with the build direction, as to minimize the overhang in the AM process. Each disk was coated with a graphite coating and placed in a testing apparatus as shown in Fig. 2. The disk was first preheated with a propane burner before it was lowered a set distance, d, into a crucible of molten aluminium A356 at temperature T for a duration of t seconds. A vacuum was applied to the chamber above the disk with a pressure of P mbar as the disk was lowered into the molten aluminium. The oxide layer on the top surface of the crucible was removed immediately before the disk was lowered into the melt pool. The test parameters are summarized in Table 2. The casting temperature and vacuum was controlled by a LPDC machine, and the disk temperature was measured by a thermocouple placed in the center of the disk" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002378_978-3-030-00036-3_5-Figure5.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002378_978-3-030-00036-3_5-Figure5.4-1.png", + "caption": "Fig. 5.4 Cube specimen for dimensional accuracy", + "texts": [], + "surrounding_texts": [ + "The actual experimentation is carried out using actual setup on FDM machine (Fig. 5.5a) (pramaanMini)manufactured bymartinjn elserman.During experimentation, the voltage of 100\u2013240 V and current of 4 A are supplied to the FDMmachine. The FDM tool head contains two nozzles: one for building of the prototype and other for building of support structure as shown in Fig. 5.5b. Filament with 1.75 mm diameter is used for fabrication of the part layer by layer. Layer height, shell thickness, and orientation angle are considered as input parameters with three levels each (Table 5.1), and ultimate tensile strength, dimensional accuracy, process time are output parameters (Table 5.2). Based on the Taguchi (L9) orthogonal array, nine experiments have been conducted to manufacture the part by FDM. Simultaneously, response parameters like ultimate tensile strength (UTS), dimensional accuracy (DA), process time (PT) are determined for each of the experimental settings. Similarly, universal testingmachine (UTM) with the capacity of 20 tons is used for determination of ultimate tensile strength for each of the nine specimens. Also, the dimensional accuracy is determined by digital caliper by measuring specimen dimensions in the directions of X-, Y-, and Z-axes. The experimental results of FDM process are shown in Table 5.2." + ] + }, + { + "image_filename": "designv11_92_0000752_sice.2008.4655087-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000752_sice.2008.4655087-Figure7-1.png", + "caption": "Fig. 7 Definitions of the stride length and the step length.", + "texts": [ + " In the \u201cpre-adaptation stage\u201d, the treadmill was in the tied configuration. In the \u201cadaptation stage\u201d, the treadmill was in the splitbelt configuration. In the \u201cpost-adaptation stage\u201d, the treadmill was in the tied configuration again. The period of each testing stage was 15 s. Treadmill acceleration and deceleration rate in switching to the next stage was 1.0m/s2. - 2510 - There are four indexesmeasured while treadmill walking in this study. The definitions of the stride length and the step length are shown in Fig. 7. The stride length is defined as the distance traveled by the ankle joint from the time of landing to the time of lift of another leg. The step length is defined as the distance between positions of the ankle joints of swing and stance legs at the time of landing of the swing leg. The duty ratio of a leg is defined as the percentage of the stance period of the leg within the walking cyclic period. The ratio of the double legs stance period (RDLSP) is defined as the percentage of the DLSP within the walking cyclic period" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000649_13506501jet353-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000649_13506501jet353-Figure4-1.png", + "caption": "Fig. 4 Test set-up for component and model tests", + "texts": [ + " The computer-based control system dynamically records the reaction force and torque, the wear and the component temperature for each single shift. There are many outside parameters e.g. wear on the locking teeth beyond the friction pair itself that influence the complex real synchronizer systems. For that reason, the system is reduced to a simple single-cone clutch, in the following referred to as \u2019component test\u2019. With a special system of adapters, also pin-on-disc type \u2018model\u2019 tests can be run on the same test rig. Figure 4 describes both test set-ups. Proc. IMechE Vol. 222 Part J: J. Engineering Tribology JET353 \u00a9 IMechE 2008 at UNIV CALIFORNIA SANTA BARBARA on July 16, 2015pij.sagepub.comDownloaded from To ensure that component tests (index K ) are comparable to model tests (index P) using plane pin-ondisc type contact it is necessary to apply the same stress to both friction systems. Therefore the average nominal pressure, pm (1), the initial sliding speed, vmax (2), and the sliding time, tr (3), ought to be equal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003050_jae-171190-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003050_jae-171190-Figure9-1.png", + "caption": "Fig. 9. Relationship between rotation radius of magnet and spring.", + "texts": [ + " linear motion part are not displaced according to the rotation angle, and momentarily linearly move at the moment the magnetic force changes. Figure\u00a08 shows the relationship between the rotation angle and the torque of the central magnets obtained. The maximum of the torque was the smallest for single-sided 4-pole magnets. This is because the angle of the magnet used for the 4-pole magnets is the largest, and as seen from the change in polarity in Fig.\u00a03, the change in magnetic flux density is considered to be gentle. Figure\u00a09 shows the arrangement of the central magnet and springs. Figure\u00a09(a) is a schematic view of the whole, and Fig.\u00a09(b) is an enlarged view of the spring attachment end of the rotating magnet. The extension length X of the spring is obtained with the rotation angle of the magnet as \ud835\udf03. One end of the tension spring is fixed to a connecting shaft at a position A of 29\u00a0mm from the magnet rotation center O. The other end is fixed at a position B of 60\u00a0mm laterally and 55\u00a0mm vertically from the magnet rotation center. The extension length of the spring is calculated from the angle \ud835\udf03. In addition, the torque T for each rotation angle obtained by the simulation was converted into the force F generated at the position A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001858_s12650-018-0495-1-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001858_s12650-018-0495-1-Figure1-1.png", + "caption": "Fig. 1 The characterization of granular flow in a horizontal rotary tumbler as determined and visualized within a past study (Romero 2012). a On top displays the six characteristic flow fields and b on the bottom displays a non-dimensional mode map of the flow fields", + "texts": [ + "1007/s12650-018-0495-1 The flow field can be characterized in a variety of ways with common non-dimensional values provided through the Froude number, in this case defined as Fo \u00bc X2D g ; \u00f01\u00de where X is the rotational rate of the tumbler, D is the tumbler diameter and g is gravity (9.81 m/s2). The Froude number represents the ratio of rotational force to gravitational force or rotational energy to potential energy. When characterized in this way, six distinct flow regimes or modes that describe the translational velocity field can be identified. Figure 1a presents an illustration of the modes, while Fig. 1b presents the modes as a function of percent fill and Froude number (Romero 2012). The exact location of mode transition can change as the percent fill of the tumbler varies, as well as the ratio between a cylindrical tumbler and particle diameters (Mardiansyah et al. 2016). This means that the Froude number alone is not sufficient to describe all possible modes of operation. This becomes particularly clear when examining variations in drum type, which may be used to generate different flow regimes more appropriate to specific commercial applications", + " Given these results, it is expected that the NSR rotation rate, x, would be a pre-definable upper bound for particle rotation (bearing in mind that particles rolling down a free surface may, on occasion, exceed this bound). Then, the lower bound would be the rotation rate of the tumbler, X, where the particle exhibits an NSNR behavior when observed by a fixed camera in the laboratory frame of reference. 3 Results and discussion 3.1 Characteristic modes The operational parameters in Table 2 were selected to achieve the six characteristic flow modes introduced in Fig. 1. Figure 5 presents photographs of the tumbler with particles in each mode of operation, showing the location of individual modes as a function of Froude number and percent fill, i.e., the mode map. These do not completely match the mode maps observed in past studies, which had been formed through data collection with significantly larger quantities of granular media. This could be a result of the rather large ratio between particle and tumbler diameters. Even with the vastly different sizes, particle density, rotation rates of the tumbler and particle system investigated here, each operational mode was achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001529_s12008-018-0471-y-Figure13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001529_s12008-018-0471-y-Figure13-1.png", + "caption": "Fig. 13 The moving direction of the contact point", + "texts": [ + " Therefore, the wear direction of the contact point should also be moved along the direction. But for boundary contact point of the liner, the wear along the normal direction mobile does not conform to the reality. According to the contact between the liner and the inner ring, the contact points that need to be updated are divided into the boundary contact points (left and right ends of liner layer) and the internal contact points except for the boundary contact point of the inner layer of the liner. The change of the contact point is shown in Fig. 13. First, the position of the contact point A in the process j th wear increment can be read from the finite element model, that is, x j and y j , these twoparameters are known, and thus r j can be calculated according to the right angle triangle AHO. r j = \u221a x2j + y2j (13) The wear depth of contact point A in the normal direction, h j , can be calculated according to the Archard wear formula, so the diameter of the inner surface of the outer ring r \u2032 j = h j + r j can be calculated after wear. Thereby the Y coordinate of the point G, y\u2032 j , can be calculated according to the right-angle triangle GHO" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.19-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.19-1.png", + "caption": "Fig. 18.19 Belt drive involving pulleys of different radii", + "texts": [ + " Calculate the total dissipation E1 at this interface and the corresponding dissipation E2 at the inner race/roller interface. (v) Finally, estimate the effective coefficient of friction for the bearing, defined as f \u2217 = (E1 + E2) (\u03a91 + \u03a92)RN . This is the coefficient of friction in a plane journal bearing of radius R which would lead to the same dissipation of energy for the same relative rotational speed between the shaft and the housing. 2. Repeat the analysis of Sect. 18.2 for the case where an elastic belt is used as a speed changing mechanism between two rigid pulleys of different radii, R1, R2 as shown in Fig. 18.19. In particular, determine:- (i) the angular velocity \u03a92 of the driven pulley, if the angular velocity \u03a91 of the driving pulley and the input torque M1 are given. (ii) what percentage of the input power M1\u03a91 is lost due to frictional slip. Which of the following parameters influence the energy loss:- \u2022 the initial belt tension, T0, \u2022 the stiffness of the belt, k, \u2022 the coefficient of friction, f , \u2022 the distance between pulley centres, L . How should we choose these parameters to minimize the power loss" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000627_isam.2009.5376920-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000627_isam.2009.5376920-Figure1-1.png", + "caption": "Fig. 1. Hardware configuration", + "texts": [ + " [1], [2] designed assembly system for tooling, fixturing and controlling. James et al. [3] also proposed analytical models suitable for each condition assembling pin-holes of various materials. Fabien et al. [4] studied assembly between high strength materials and assembly forces depending on the shape of pin section. But none have focused on repeated insertion. The present research focuses on changes caused by repeated insertion, and analyzed the relationship between interference fits and forces of repeated insertion. II . EXPERIMENT Figure 1 and Table 1 show the hardware and its specifications. An aluminum plate with holes was fixed to a dynamometer, and the dynamometer was fixed to an X-V-stage. A pin was fixed to a Z-stage and the pin-hole insertion occurred in the Z-direction. Figure 2 illustrates the process of fabricating the hole and the process of measuring the insertion force. The holes were fabricated with a drill and an end mill. 978-1-4244-4628-5/09/$25.00 \u00a92009 IEEE 107 Two pins were made from tungsten carbide (WC-Co, AF310, Axismateria, Japan)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000746_cso.2009.275-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000746_cso.2009.275-Figure1-1.png", + "caption": "Figure 1. Schematic of sliding contact", + "texts": [ + " The standard linear solid model and Burgers model can\u2019t reflect the influence of frequency; the generalized fractional derivative model and four-parameter model is able to describe the influence of frequency exactly, but the calculation formula was complex; Liu [5] established the fractional derivative Kelvin model and Maxwell model and analyzed the variation rules of storage modulus and loss modulus with the frequency. However, the above models didn\u2019t take both the temperature and frequency into account. Because friction lining is a kind of polymer, its mechanical property is strongly dependent on the temperature and loading frequency [6]. In order to master the thermo-stress coupled behavior, it is necessary to describe the thermoviscoelastic constitutive relation of friction lining quantitatively. Figure 1 shows the schematic of sliding contact between friction lining and wire rope. From Figure 1, it is seen that the friction lining contact with the out strands of wire rope. The lining is subjected to a periodical load. And there is rapid temperature rise of friction lining in a short time during high-speed friction sliding with wire rope [7]. From the above analysis, it is obvious that the friction lining is subject to temperature and frequent load during the process of high-speed sliding and their effect can\u2019t be ignored. The generalized Kelvin model was used in this study and its constitutive equation is defined by F Q\u03c3 \u03b5= , (1) 978-0-7695-3605-7/09 $25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003979_robomech.2019.8704789-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003979_robomech.2019.8704789-Figure1-1.png", + "caption": "Fig. 1. Two-link wheeled mobile manipulator", + "texts": [ + " Using the optimality condition for controls \u2014\u2014 = 0 and solving for optimal ou velocities yields $2 , m = Arep Gxp Arep Gxp 1 f ( p j j 2 \\ 21 J 1 ( { p j m ) 2 \\ 2 I ) if 0 < 7 < 1 and Pi,m < f i b s , if 7 > 1 and p\\ n < u 1 u2 U3 \u2014(Ai cos ip + X2 sin cot \u03b1. A subclass of \u201cidealized\u201d plates was introduced in [4], for which the equality K(\u03b1) \u2261 cot \u03b1 is fulfilled in the range of \u03b1\u0303 < \u03b1 < \u03c0 \u2212 \u03b1\u0303 (Fig. 2). In this paper we mainly consider the case where the blade is an idealized plate with a high aerodynamic quality. Now, when the phenomenological model of the object is fully defined, we can proceed to constructing a mathematical model. The dynamics of a rotational body in the problem under consideration can be described by a system of differential equations: u\u0307 = b[CL(\u03b1) cos \u03b8 \u2212 CD(\u03b1) sin \u03b8]w2, v\u0307 = \u22121 + [CL(\u03b1) sin \u03b8 + CD(\u03b1) cos \u03b8]w2, (2.1) where u = w sin \u03b8 and v = \u2212w cos \u03b8 are dimensionless rotational speed of the blade (point C) and the speed of translational motion, w2=u2+v2; \u03b8=\u2212 arctan(u/v), if v < 0 and \u03b8 = \u2212 arctan(u/v) + \u03c0 sign u, if v > 0; \u03b1 = \u03d5 \u2212 \u03b8; dot on top denotes the derivative with respect to the dimensionless time \u03c4 , and u = \u03a9r V\u2217 , v = V V\u2217 , \u03c4 = g V\u2217 t, V 2 \u2217 = 2mg \u03c1S , b = mr2 J ", + "4) We note some general properties of the solutions of this equation in the case of an idealized blade with HAQ [4]. a) For \u03d5 = 0 equation (3.4) has a unique solution \u03b1\u0304 = 0, and when \u03d5 = \u03c0 \u2013 a unique solution of \u03b1\u0304 = \u03c0. b) If \u03b1\u0304 is a solution of (3.4) for some value of \u03d5 different from \u03c0/2, then for a setting angle \u2212\u03d5 there exists a solution \u03c0 \u2212 \u03b1\u0304. Therefore, only positive values of a setting angle will be considered in the following. c) A special situation occurs for \u03d5 = \u03c0/2. In this case, there exist a pair of isolated solutions \u03b1\u03041 and \u03b1\u03042 = \u03c0 \u2212 \u03b1\u03041 and a range [\u03b1\u0303, \u03c0 \u2212 \u03b1\u0303] of non-isolated solutions (Fig. 2). MECHANICS OF SOLIDS Vol. 53 No. 1 2018 3.1. Constructing the Dependence \u03b1\u0304(\u03d5) We now consider dependence of solution of (3.4) on different values of angle \u03d5. Fig. 3 shows a qualitative diagram of the graphical solution of this equation for an idealized plate with HAQ. Thick line denotes the graph of K(\u03b1), the fine lines represent the graphs of the function tan(\u03d5 \u2212 \u03b1) for certain values of \u03d5. Let \u03d5 be somewhat bigger than \u03c0/2. Obviously, both isolated solutions will be displaced somewhat, and the value of \u03b1 increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.4-1.png", + "caption": "FIGURE 16.4 Plasma formation in keyhole welding.", + "texts": [ + " This results in hightemperature gradients in the immediate vicinity of the heat source, producing highcooling rates once the beam is switched off. Thus, Q-switching (see Section 6.2) is not normally used for deep penetration laser welding because the shorter pulse duration of the Q-switched laser results in lower penetration. It is more appropriate for microwelding. 16.1.3.1 Plasma Formation With high-power beams, a portion of the laser energy is expended in ionizing part of the shielding gas and/or vapor cloud that forms above the workpiece, to form a plasma (Fig. 16.4). This plasma may absorb part of the laser beam energy and reradiate it in all directions or scatter it, thereby reducing the amount of energy reaching the molten pool. The plasma cloud may also induce an optical lensing effect that changes the effective focusing of the laser beam. This laser-induced plasma normally develops in the intensity range of 106 W/cm2 when welding steel with CO2 lasers. Even though initial absorption of a CO2 laser beam on a metallic surface is normally less than 10%, the absorption can be as high as 100% once a plasma is formed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001724_icit.2018.8352228-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001724_icit.2018.8352228-Figure4-1.png", + "caption": "Fig. 4 Image of the test rotor supported by oil\u2013film bearings, which is rotated and radially positioned by the consequent-pole-type BELM.", + "texts": [ + "05 mm in order to verify the independency of from the eccentricity. Fig. 3 (b) shows the simulation results. The force current coefficient is proportional to the eccentricity in the eccentricity range from 0 to 0.2 mm. This means that the coefficient of is almost a constant value. The slope is approximately = 18.5 N/A mm. In a similar manner to the principle shown in Eq. (4), the simulation results represent that the radial excitation force is proportional to the eccentricity and the d-axis current. III. EXPERIMENTAL EVALUATION Fig. 4 shows an image of the test rig, and Table 2 lists its specifications. The rotor of the test rig was supported by an oil\u2013 film bearing at either end. The rotor displacement was measured using two pairs of eddy\u2013current\u2013type displacement sensors (PU\u201305, Applied Electronics Corp.). The outputs of the displacement sensors placed on either side of the motor were averaged to measure the displacement at the center of the motor. At the right edge, there is an encoder to detect an absolute rotational angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002974_3284179.3284195-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002974_3284179.3284195-Figure1-1.png", + "caption": "Figure 1. Simple pendulum", + "texts": [ + " In this paper we propose the use of VidAnalysis free app [25] to effortlessly track the position of the each of the spheres using a smartphone. This intuitive and easy-to-use app allows the students to track the position of the spheres in their smartphones, through screen touching, frame by frame, generating position versus time graphs. The data can then be exported into a .csv file and manipulated, either directly in the smartphone or using a computer. 2.1.1 Simple pendulum (one sphere). Two forces act on a bob of a simple pendulum: the tension of the string and the gravitational force, as shown in Figure 1. Permission to make digital or hard copies of all or part of this work for personal or classroom use is granted without fee provided that copies are not made or distributed for profit or commercial advantage and that copies bear this notice and the full citation on the first page. Copyrights for components of this work owned by others than the author(s) must be honored. Abstracting with credit is permitted. To copy otherwise, or republish, to post on servers or to redistribute to lists, requires prior specific permission and/or a fee" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001077_caidcd.2008.4730527-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001077_caidcd.2008.4730527-Figure1-1.png", + "caption": "Fig 1. Sketch of a rolling cut shear l-erank; 2-link rod; 3-upper blade carriage; 4-lower blade carriage; 5-steel plate; 6-oriented rod; r-erank radius;", + "texts": [], + "surrounding_texts": [ + "Keywords: Rolling cut shear, Steel plate production, Shearing quality, Strength parameters, CAD\n1. Introduction\nRecently, iron and steel industry of China has got great development. The production of iron and steel was over 400 million tons in 2006. More production lines of steel plate have been set up. In accordance with the developments, the demand for equipment of steel plate production is greatly increasing and the requirement for the quality and performance of shearing machines is getting higher.\nDue to the advantages such as good shearing quality, higher shearing efficiency, lower wearing of blade, smaller distance of stroke and lower peak value of shearing force, rolling cut shears are taking the place of traditional inclined throat shears as the main shearing equipments of steel plate production[1-4]. With the increasing demand of users for the quality of products, the requirements of steel rolling plants for shearing quality and performance of rolling cut shears are getting higher and more requirements for the design and manufacture of rolling cut shears are also put forward. The ideal working state of a rolling cut shear is pure rolling cut. For the reason of technique, it's hard to\n978-1-4244-3291-2/08/$25.00 \u00a92008 IEEE\nrealize pure rolling cut. Only nearly pure rolling cut can be obtained for the widely used link-rod type rolling cut shears at present. Strength parameters of rolling cut shears such as shearing force and drive moment etc. and the load of mechanism are closely relative with the design of the machine[5-7]. For the optimization of mechanism design, a CAD software of rolling cut shears has been developed and was used for analysis of those factors such as crank radius, link rod length, phase difference, the center distance of cranks and the offset position of upper blade carriage, which influence the shearing quality and strength parameters, under given conditions (plate width 3500mm, plate thickness 40mm, ultimate strength 700MPa).\ne-the offset position of upper blade carriage; L-link rod length; B-width of steel plate\n2. Influence analysis of design factors on shearing quality\nIn the ideal state of pure rolling cut, the motion track of the lowest point of upper blade should be a straight line and the shearing quality be the best. There are so many factors that influence shearing quality of rolling cut shears. The Influences of crank radius, link rod", + "length, phase difference and the center distance of cranks are analyzed below.\nIt can be seen from fig.2 that the influence of crank radius on shearing quality is significant. When crank radius is 110mm, the difference of the motion track of the lowest point of upper blade is only 1.32mm. That is nearly pure rolling cut and shearing quality is the best. With the increase of crank radius, shearing quality is becoming worse.\n] 6\n.\u20ac 4 1.r 2\ni 0 100 105 110 115 120 125 130 135\ncrank radius/mm\n2.96mm to 2.36mm. The variation is only 0.6mm. So, if the method of changing link rod length is adopted, other factors should be synthetically considered.\nThe influence of phase difference on shearing quality is great. When the phase difference is 55\u00b0 and 60\u00b0, the difference of the motion track of the lowest point of upper blade is only 2.46mm and 2.7mm respectively and these values are smaller. No matter the phase difference is increasing or decreasing, shearing quality will rapidly become worse.\nThe influence of the center distance of cranks on shearing quality is also significant. Within analytical range of 2000mm to 4000mm, when the center distance of cranks is 3250mm, the difference of the motion track of the lowest point of upper blade is 1.98mm. At that point, shearing quality is the best. No matter the center distance of cranks is increasing or decreasing, shearing quality will become worse.\nthe center distance of cranks/mm\nStrength parameters are main basis of mechanical strength design and the choice of driving motor power. For a rolling cut shear, main basis of design is shearing force and resultant moment. The peak value of shearing force is usually much greater than normal value. It is the main factor influencing mechanism strength. So, it is the responsibility of designers to reduce the peak value of shearing force as possible by choosing reasonable design parameters. There are many factors that influence the strength parameters. The attention of this paper is mainly put on those such as phase difference, crank radius, link rod length, the offset position ofupper blade carriage and the center distance of cranks etc.\nIt can be seen from fig.3 that the influence of phase difference on strength parameters of rolling cut shears is significant. While phase difference increases from 45\u00b0 to 75 0, the peak value of shearing force decreases from 16881 KN to 8168KN and the resultant moment decreases from 1502KN-m to 723KN-m. Both variation ranges exceed 50%. The reason is that the bigger the phase difference is, the greater the equivalent cutting angle will be at beginning cutting stage. That makes cutting area as well as the peak value of shearing force and resultant moment reduced. In this sense, it seems that the bigger the phase difference is, the more benefit there will be for reducing strength parameters. But we have known from above analysis that as phase\n] 3 i 2.8 2.6\ncrt 2.4 \u00b7r 2.2 ~ 2 00\n800 850 900 950 1000\nlink rod lengthlmm\n] 20 .\u00a3 15 1 10 OJ)\n5 .~ ..t:: 0 00\n45 50 55 60 65 70 75\nphase difference/degree\nS 20,:~ t=:'E?~~~ i 2000 2500 3000 3500 4000\nFig 2. Influences of design factors on shearing quality\nThe shearing quality is getting better with the increase of link rod length, but the influence is not so prominent as a whole. While link rod length increases from 800mm to 1000mm, the difference of the motion track of the lowest point of upper blade varies from" + ] + }, + { + "image_filename": "designv11_92_0000035_cp:20080207-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000035_cp:20080207-Figure1-1.png", + "caption": "Figure 1: Zoom-in on one pole pair of a typical 2D vector potential plot in an induction machine with the corresponding skeleton (white). The surfaces associated with fluxes ~k are indicated.", + "texts": [ + " From the point of view of Thermodynamics, the balance of magnetic energy writes (2) The rate of magnetic work I / V - / f l j ' \u00a3 ~ a - fa ~ n x ha\" \u00a3v a (3) nent geometrical structure when the machine is operated in steady state operation. This regularity is a redundancy one attemps to get rid of by defining the equivalent circuit. As the equivalent circuit counts a small number of lumped parameters, identification is in general done empirically, as reported in the abundant literature on the subject [2]... [8], and no general (theoretical) definition of the lumped element of the equivalent circuit is available. As mentioned above, all vector potential plots look more or less the same, Fig. 1. As the vector potential (a differ- ential 1-form) evaluates on curves [1], the idea is to select the characteristic (closed) curves whose associated fluxes give most information about the field. In electrical machines, due to the symmetry, one can restrict the analysis to one,pole pair of the machine and it is enough to select the 3 curves C1, C2 and C that enclose the maximum fluxes in the stator the rotor and the air gap regions respectively. Those fluxes are accordingly noted ~1, ~2 and ~ and represented in Fig. 1 and Fig. 2. The 3 characteristic curves define the cylinder-shaped topological structure depicted in Fig. 2, and which can be considered as the skeleton of the a field. This skeleton is fixed with respect to the rotating field, i.e. it rotates at the speed \u20220 = 27fro/p, where fo is the frequency and p the number of pole pairs. By Poincare duality, the skeleton is associated a dual topological structure, the co-skeleton, which has the shape of a pretzel, Fig. 2. Whereas the characteristic fluxes ~k in the machine are associated with the closed curves of the skeleton, the characteristic magnetomotive forces fk and currents ik are associated with the curves (not necessarily closed) and the surfaces of the co-skeleton" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000569_isma.2008.4648827-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000569_isma.2008.4648827-Figure6-1.png", + "caption": "Figure 6. Pendulum energies", + "texts": [ + " Figure 5 is the root-locus diagram of the plant poles when changing the proportional gain. Based on the selection of the proportional gain, the system is either unstable or critically stable. If the gain is selected to result in a critically stable system, the system is critically stable in the reaching mode. However, for the entire system controlled using the proposed SMC, the following simulation will show that the system is asymptotically stable due to the sliding action along the sliding surface. With reference to Figure 6, in order to find the optimal limits of Equations 22 and 33, the pendulum is disturbed by applying an initial input angle\u03b8 . The pendulum in this position has a maximum potential energy Q that must be compensated for by the kinetic energy H in order to keep the system stable. The kinetic energy H should be attained from the cart motion. To guarantee this, the following equation is used: It should be noted that Equation 24 is valid under the assumption made earlier that the pendulum angular velocity is negligible. For small input angle, 1cos =\u03b8 and Equation 24 can be cast in the following form: mM mglx + = 2 (25) Substituting for the values of the parameters from Table 1 into Equation 25 gives sec/683.0 mx = . With reference to Figure 6, assume the following condition is valid: xl =\u03b8 (26) Substituting for ml 5.0= and sec/683.0 mx = into Equation 26 gives sec/366.1 rad=\u03b8 . Assuming the system state is on the sliding surface of the pendulum, for an initial input of rad1.0\u2212=\u03b8 , the slope of the sliding surface can be approximated to be: 1 2 sec66.13 1.0 336.1 \u2212= = = \u03b8 \u03b8\u03bb (27) The initial state of the inverted pendulum system is boundary stable upright. The system is disturbed by applying an initial input of an angle of rad1.0\u2212=\u03b8 ( O73" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002915_978-3-030-04792-4_31-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002915_978-3-030-04792-4_31-Figure3-1.png", + "caption": "Fig. 3. An additional proximity sensor is used as a keyphasor.", + "texts": [ + " In order to measure the relative vibration between the shaft and bearing housing, two proximity probes of 90\u00b0 phase angle are installed at each bearing housing in the horizontal (named as P1X or P2X) and vertical (named as P1Y or P2Y) direction (see Fig. 2). The specifications of these sensors are listed in Table 2. Table 1. Specifications of the test rig. Shaft Material Steel Diameter (mm) 10 Length (mm) 580 Young modulus (GPa) 205 Density (kg/m3) 7800 Mass (kg) 0.5 Disk Material Steel Diameter (mm) 75 Thickness (mm) 25 Mass (kg) 0.5 Bearing Mass (kg) 0.150 Stiffness (horizontal direction) (N/m) 1.25 105 Stiffness (vertical direction) (N/m) 3.83 108 As mentioned before, an additional proximity sensor is used as a keyphasor (Fig. 3). This keyphasor is very important and useful for converting the signal from the time domain to angular domain for performing the \u201cTacho\u201d signal, synchronous averaging, motor speed, balance of the system and so on [11]. Figure 4 shows the signal acquired from two proximity sensors at the bearing #1. The signal from bearing #2 is quite similar and not shown here. It can be seen that the amplitude of these sensors (or the relative vibration between the shaft and bearing housing) increases dramatically at about 8th second from the beginning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000881_pi-b-2.1962.0229-Figure21-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000881_pi-b-2.1962.0229-Figure21-1.png", + "caption": "Fig. 21.\u2014Positions of i.l.s. glidepath slot aerial. [a) Optimum position for aerial in metal nose.", + "texts": [ + " and an omnidirectional aerial for the same equipment. The slot has also been used as an i.l.s. glidepath receiving aerial in the nose of an aircraft. A particular virtue of the slot aerial is that, if symmetrically mounted in a conducting surface, it will have a symmetrical radiation pattern, even if fed '6) Position for aerial behind radome. from a coaxial cable. It has been found that this type of aerial is less susceptible to radiation pattern variations due to the movement of search radar scanners adjacent to the aerial. Fig. 21 (a) shows the optimum position for the aerial in a metal nose. Where a radome is fitted, the aerial can be mounted in the underside of the nose behind the radome [Fig. 21(6)], provided that the chord through the ends of the slot subtends an angle with the horizontal greater than 20\u00b0. At smaller angles the signal in the horizontal plane is too weak. (7) AERIALS FOR AUTOMATIC LANDING EQUIPMENT (7.1) General Requirements The radio aids developed by the Blind Landing Experimental Unit (B.L.E.U.) of R.A.E. have been described by Shayler.5 The requirements of this system necessitate extra care in the choice and siting of the aerial systems, and some of the precautions are described below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003691_robio.2018.8665156-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003691_robio.2018.8665156-Figure4-1.png", + "caption": "Fig. 4. Configuration of axes on board [14].", + "texts": [ + " After acquiring the image data every 1 ms, the image-processing PC measured the position and orientation of the board within 1 ms and sent the measurement results to the real-time controller via Ethernet using the UDP protocol. The board had a length of 220 mm, a width of 100 mm, a thickness of 5 mm, and a mass of about 136 g. Retroreflective markers were attached at four corners of the board to simplify corner detection by the high-speed camera. The configuration of axes on the board is as shown in Fig. 4. Here we explain the overall strategy for achieving the human\u2013robot collaboration. This section gives a brief explanation of the proposed strategy. The details of the proposed strategy are given in [14]. We show the control flow in Fig. 5. We divided the human\u2013 robot collaboration into the following two steps: 1) The position and orientation of the board are measured using high-speed image processing (Section III-B). 2) The robot hand is controlled according to the position and the orientation of the board (section III-C)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003206_s0869864318050104-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003206_s0869864318050104-Figure4-1.png", + "caption": "Fig. 4. Diagram of the cladding of overlapping beads.", + "texts": [ + " To solve the system of equations (10) and (11) it is necessary to specify the angle 1i\u03b3 \u2212 characterizing the side point of the beads contact, which may be found from the equation ( ) ( ) ( )( ) 1 210 0 cos 2 . 2 1 cos i i i d b z \u03b3 \u03be \u03be \u03be \u2212 \u2212 = + \u2212 \u222b (12) 1.4. Cladding of overlapping beads At the laser cladding for obtaining a continuous coating the case occurs frequently when the beads are formed by an overlapping cladding. In such a situation, the beam is shifted to the right by the value d (Fig. 4), which is equal to the distance between the centers of beads 1iO \u2212 and .iO The value of overlapping \u2206 is uniquely determined if the widths of the both beads 1 andi ib b\u2212 and the shift d are known, then ( )1 2 ,i ib b d\u2212\u2206 = + \u2212 and .ib\u2206 < The method proposed in the work [13] for constructing the profiles of the beads formed by overlapping cladding is based on a local replacement of the curve curvilinear part connecting the points M and N by a straight-line segment MN (Fig. 4), which reduces the accuracy of computations. Consider the algorithm where this inaccuracy has been eliminated. The system of equations for determining the unknown parameters 0 1 2, ,i i iz \u03b8 \u03b8 has the form ( ) ( )( ) ( ) ( )2 0 1 1 22 1 cos sin sini i i i i bf b z \u03b8 \u03b8 \u03b8= + \u2212 \u2212 + \u2212 ( ) ( )( )21 1 1 1 0 1 12 1 cos sin sin ,i i i iz \u03b8 \u03b8 \u03b3\u2212 \u2212 \u2212 \u2212 \u2212 \u2206 + \u2212 \u2212 + (13) ( ) ( ) ( ) ( )2 21 1 1 1 0 1 02 1 cos 2 1 cosi i i iz z\u03b8 \u03b3\u2212 \u2212 \u2212 \u2212+ \u2212 \u2212 + \u2212 = ( ) ( ) ( ) ( )2 2 0 1 0 22 1 cos 2 1 cos ,i i i iz z\u03b8 \u03b8= + \u2212 \u2212 + \u2212 (14) ( ) ( ) ( )( ) 2 1 2 0 cos . 2 1 cos i i i i d b z \u03b8 \u03b8 \u03be \u03be \u03be = + \u2212 \u222b (15) Equation (13) describes by analogy with (7) or (10) the balance of forces acting on the fluid. Equation (14) has been obtained from the condition of the equality of the contact point coordinate c (Fig. 4), which is calculated for the left ( 1)i \u2212 -bead as ( ) ( )21 1 0 12 1 cosi ic z \u03b8\u2212 \u2212= + \u2212 \u2212 ( ) ( )21 1 0 2 1 cosi iz \u03b3\u2212 \u2212+ \u2212 and for the right i-bead as ( ) ( )2 0 12 1 cosi ic z \u03b8= + \u2212 \u2212 ( ) ( )2 0 22 1 cosi iz \u03b8+ \u2212 . Equation (15) yields a relation between the width of the i-bead ib and its limiting contact angles 1 2,i i\u03b8 \u03b8 at contact points that are measured from the horizontal, which closes the system of nonlinear equations. The value of the angle \u03b3 i\u20131, as (11) above, is determined using the width of the (i \u20131)th bead: ( ) ( ) ( )( ) 1 1 210 0 cos 2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001674_978-3-319-76138-1_3-Figure3.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001674_978-3-319-76138-1_3-Figure3.2-1.png", + "caption": "Fig. 3.2 Statics of a cable robot of type 2T with four cables and 3T with purely translational motion pattern", + "texts": [ + " Since this cannot be fulfilled exactly in practice, parasitic motion can occur that is neither described nor postulated with the specific models. Anyway, these models can be very useful as approximation or as model system. All motion patterns other than 3R3T can only be achieved by non-generic geometries, i.e. special constraints must be fulfilled for the geometric parameters. In a certain sense, all these robots are architecturally singular designs that lead to a very specific degeneration of the robot\u2019s motion pattern. Firstly, the planar and spatial case for a point mass is addressed where no rotation motion occurs (Fig. 3.2). To set up a planar or spatial robot without rotational motion, the cables have tomeet in a common point on the platform.Without loss of generality, we can assume that all vectors bi are equal and we can translate the platform\u2019s reference point such that all bi = 0. Thus, we consider a pure force equilibrium and the structure matrix becomes AT = [ u1 . . . um ] , (3.10) where the vectorsui = [ui,x , ui,y]T \u2208 IR2 for the 2T case andui = [ui,x , ui,y, ui,z]T \u2208 IR3 for the 3T case. The applied wrench becomes wP = [ fx , fy]T \u2208 IR2 and wP = [ fx , fy, fz]T \u2208 IR3, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002955_978-3-030-03451-1_54-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002955_978-3-030-03451-1_54-Figure2-1.png", + "caption": "Fig. 2. a. Static unbalance caused by eccentric centre of mass S\u2019, b. Rotor consisting of a set of static unbalances, c. Influence of a static unbalance ~Uk on the resulting unbalances ~UA and ~UB (amended from [2]).", + "texts": [ + " This section aims to give an overview over unbalances that may occur through deviations during the manufacturing or assembly of permanent magnet rotors. Static Unbalance. The basic type of an unbalance is the case of a static unbalance. The unbalance vector (unbalance magnitude and phase) can be described as the result of an eccentric mass distribution within a rotating object with relatively low axial expansion [2]. A commonly used unit for the description of a static unbalance is \u201cgmm\u201d and can be seen as the product of an unbalance mass and its distance to the axis of rotation (Fig. 2a). This type of unbalance can be detected either through rotation and vibration analysis or the use of a focal point scale without rotation [2]. Couple Unbalance. Another specific case of a rotor unbalance is a couple unbalance. It is caused by two complementary unbalance vectors laying in two separate planes, having the same magnitude and pointing in opposite directions. This unbalance can only be detected by rotating the rotor [2]. Dynamic Unbalance. In practice, the unbalance state of a rotor consists of a combination of static und couple unbalances (Fig. 2b). In the case of a rotor that consists of a shaft and several disc elements, the resulting unbalance can be calculated through the combination of the respective unbalance vectors of each disc (Fig. 2c). This combined unbalance state is called dynamic unbalance and is commonly modelled by two unbalance vectors laying in separate planes but pointing in two independent directions and/or with different magnitudes [4]. The admissible residual unbalance of a rotor is defined by the balancing grade. Taking into account the rotor\u2019s overall mass, the operating speed and the axial position of the centre of mass, a maximum admissible dynamic unbalance of a rotor can be calculated [5]. For assessing the measuring uncertainties occurring during unbalance measurement, standardized rotor designs with specified initial unbalances can be used [6]", + " Through the introduction of separate coordinate systems per part and assembly respectively position deviations can be modelled separately and used for statistical analyses. Conservation of momentum and angular momentum for each rotor component is used to predict the overall rotor unbalance. If, for instance, the unbalance contribution of a rotor disc is to be calculated, the model uses the position and angular orientation of a disc with respect to the axis of rotation together with the disc\u2019s mass and static unbalance. 3D-models and technical data of the two rotor designs that were used for this article are shown in Fig. 2 and Table 1. Optimized Assembly Strategies. Since different strategies can be chosen to assemble the shaft, lamination stacks and magnets of a permanent magnet rotor [11], several approaches for an optimized assembly are possible. In this case the focus was put on the implementation of an online optimization strategy that doesn\u2019t need a binning of rotor components before reaching an optimized arrangement. Even though online optimized assembly mostly only leads to a local optimization, it allows time efficient calculations and reduced logistic effort during production", + " The optimization process for both rotor designs can therefore be split up into the following two steps: \u2013 Online optimized assembly of magnets onto or into a single discs \u2013 Online optimized assembly of discs (with magnets) onto a shaft For each of these two optimization and mounting steps, several optimization algorithms can be applied. Algorithms such as greedy algorithms, backtracking search and metaheuristic optimization are presented in the following sections. For a first validation of the model-based approach, investigations were carried for the ProLemo rotor design (cf. Fig. 2a and Table 1). In order to analyse the effects of the different components on the rotor unbalance, separate investigations have been done on the shaft unbalance, shaft eccentricity, disc unbalance (with and without magnets mounted) and the resulting rotor unbalance. Shaft Unbalance. A set of 20 shafts was produced to analyse how the different manufacturing steps on the shaft affect its unbalance. Comparing the measurements of unbalances after the three successive production steps (turning, deep-hole drilling and toothing), significant differences in the effects on the shaft unbalance could be detected: The drilling of the deep holes resulted in an average unbalance that was ten times higher than the average unbalance of the shafts right after the turning process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001888_s2075113318030176-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001888_s2075113318030176-Figure1-1.png", + "caption": "Fig. 1. Scanning an electron beam over the substrate coated with the composite layer: (1) first stage of pumping; (2) electron beam transported to the atmosphere; (3, 4) areas illustrating the position of electron beam on the substrate; (5) scanning trace (path of electron beam motion); (6) coordinate table; (7) substrate; (8) layer of the composite powder.", + "texts": [ + " Consequently, an adiabatic temperature of interaction of titanium and chromium oxides with carbon, Tad, in the self-propagating high-temperature synthesis (SHS), which reaches 3700 K, was determined [9], and the equilibrium composition of products of heterogeneous interaction was determined. Table 2 shows the chemical compositions of the reaction mixtures used in experiments to form coatings on the surface of 12Kh18N10\u0422 steel. With such formulations of reaction mixtures, the formation of TiC, Cr3C2, and Fe3C or Cr3C2 and Fe3C carbides and the presence of titanium nitride TiN and oxide Ti2O3 are possible. Figure 1 shows the scheme of the surfacing process. Surfacing by electron beam 2 of coatings from mixtures of TiO2 : 2.1C and TiO2 : 0.3Cr2O3 : 3.3C were deposited on a substrate 7 of 12Kh18N10T steel with a surface area of 75 \u00d7 75 mm and a thickness of 8 mm by using an electron beam in the atmosphere. A layer of the composite powder 8 mixed with an organic binder was preliminarily deposited on the substrate surface. Using an industrial robot KUKA KR 240 R3200 PA, an electron beam with energy of 120 keV was mechanically swept over the substrate surface with a deposited composite layer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure28-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure28-1.png", + "caption": "Fig. 28 Cable guide", + "texts": [ + " The actuator unit by means of ball screw has the advantage of high efficiency and no backlash. Two cables actuating the knee joint are wrapped around the pulley, the axis of which coincides with that of knee joint. By virtue of this structure, the moment arm on knee joint torque is constant regardless of knee joint angle. As a consequence, the wide range of joint movement can be achieved. Each edge point of three cables actuating the ankle joints has the cylinder-shaped cable guide as shown in Fig. 28. The axis of each cylinder is parallel to the pitch axis. This wire guide prevents the cables bending below its minimum bend radius and diminishes the moment arm variation around the ankle joints. Figure 29 shows the control flow of the wire-driven biped robot. The blocks in Fig. 29a are the same as the controller described previously in Sect. 3; hence, the details of these blocks are omitted. The control procedures specific to the wiredriven robot are the blocks in Fig. 29b, which are designed to enhance the joint servo performance of wire drive system as in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002481_gt2018-76823-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002481_gt2018-76823-Figure2-1.png", + "caption": "FIGURE 2. TEST BOX SKETCH", + "texts": [ + " The gear is clamped on top of the shaft by a proper fixing system easy to be unmounted when necessary. The gears used for the experimental investigation have a face length of L f /Dp = 0.36 and 38 teeth. 3 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 05/06/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use A D j/Dp = 0.0065 diameter oil jet, provided by a spray-bar placed at a distance of L j/Dp = 0.013 was used to lubricate and cool the gears. The jet was directed radially against the driving gear. Figure 2 shows a sketch of the test box containing the gears and the spray-bar, the box is quite wide in order to reduce the volume effect on windage losses and to be as close as possible to free gear conditions. This choice was due, in a first phase, to the desire of comparing the results for the single gear configuration with preceding correlations (see [15]), and now to highlight the effect of gear meshing with respect to the preceding results. For these reasons the same box, of dimensions ( W Dp \u00b7 H Dp \u00b7 L Dp ) = 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000007_amr.44-46.829-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000007_amr.44-46.829-Figure7-1.png", + "caption": "Fig. 7 Investigation categories for present EPFE analysis", + "texts": [ + " As shown in Figure 5, this tilting bogie has a special characteristic that the force from side frame of the bogie is transferred to bearing through the wing block acting on the adapter. The wing block can wave right or left around its rotational centre with vehicle vibrating laterally. The wheel set is subjected to a dynamic variable balance force system, as shown in Figure 6. A wheel-track contact force spectrum inspected on-line on a R300 curved line is used for studying the bearing distributed fatigue stresses. With respect to the selected inspection positions shown in Figure 7, results of the EPFE analysis are exhibited in Table 3. It verifies that 1. Rollers and inner- and outer-rings of the bearing are subjected to unbalance distributed axial equivalent stress amplitudes. The closer to axle side, the larger the equivalent stress amplitude. 2. Maximum equivalent stress amplitude appears at the outer ring inboard side close to the seal seat. K6 Type Bogie. At the same time, 353130B bearings are also used on the same axle weight freight cars with K6 type bogie. Number of bearings on K6 type of bogie is actually more than those on K5 type bogie", + " An EPFE analysis is carried out on the bearing on K6 type of bogie to understand the difference of two kinds of bogie. As shown in Figure 8, this kind of bogie has a structure of down crossing bars to control the vehicle waving laterally right or left. Force from the side frame of bogie is transferred to bearing through rubber acting directly on the adapter. By the wheel-track contact force spectrum under same piece of railway line and same speed, results of EPFE analysis are displayed in Table 3 with respect to the selected inspection positions shown in Figure 7. It verifies that 1. The equivalent stress amplitudes show heterogeneously axial distributions for rollers and inner- and outer-rings of the bearing for K6 type of bogie. 2. The maximum equivalent stress amplitude is still at the outer ring inboard side close to the seal seat for K6 type of bogie. But the value of 293.855 MPa is slight lower than the value of 299.442 MPa for K5 type of bogie. 3. Distributed stress amplitudes of the rollers are obviously lower than those for K5 type of bogie. The above indicates that the lower fatigue stressing is a cause of no occurred failure for the bearings on K6 type of bogie" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002393_gt2018-75983-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002393_gt2018-75983-Figure3-1.png", + "caption": "Figure 3 Cross-section view of the rotordynamic test rig", + "texts": [ + " The orifice configuration the three different sets is summarized in Table 1. 3 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 09/17/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use Rotor consists of five main components: two bearing journals, an induction motor element, a thrust runner, and an end cap (with the same weight as the thrust runner). All the components are assembled with a set of tension bolts. Table 3 presents the overall geometries and parameters of the rotor. Figure 3 depicts the cross-section of the rotordynamic test rig. The test rig is driven by a 10 kW, two-pole, asynchronous, highspeed induction motor. The cooling jacket around the motor stator provides it with thermal management. The cooling jacket has circumferential grooves (channels) and pressurized air is injected into the grooves to remove the heat from the motor stator. The shaft is supported by two HAFBs located at each side of the motor. The axial motion of the shaft is supported by a set of air foil thrust bearings (AFTBs)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002408_978-3-319-99522-9_9-Figure9.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002408_978-3-319-99522-9_9-Figure9.1-1.png", + "caption": "Fig. 9.1 Kinematical scheme of the leg A of upside-down mechanism", + "texts": [ + " The elements of its legs have known size and given mass. To simplify the graphical image of the kinematical scheme of the mechanism, in what follows we will represent the intermediate reference systems by only two axes. The zk axis is represented for each component element Tk. It is noted that any relative rotation with angle uk;k 1 of the body Tk must always be pointed along the direction of the zk axis. One of the three active elements of the robot is the input link of the first upside-down leg A (Fig. 9.1). This is a homogenous crank of length A1A2 \u00bc l1 and mass m1, which rotates about z A1 axis with the angle uA 10, angular velocity xA 10 and angular acceleration eA10. The total tensor of inertia of the input link-rotor mounted on this link is J\u03021. The centre of the transmission rod A3A5 \u00bc l2 is denoted by A2. This link is connected to the frame A2xA2 y A 2 z A 2 (called T A 2 ) and has a relative rotation about axis z A2 with the angle uA 21, so that xA 21 \u00bc _uA 21 and eA21 \u00bc \u20acuA 21. It has the mass m2 and the central tensor of inertia J\u03022" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000622_icinfa.2009.5205077-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000622_icinfa.2009.5205077-Figure6-1.png", + "caption": "Fig. 6. Geometric diagram.", + "texts": [ + " Then a terrain inclination map { ( )}M M M dh x needs to be extracted as a reference for slope measurement at the point 2C . The variable dx represents the location of the robot on the path. In this example, the whole map is divided into eight sections. The slope function is 0( ) 0M d Mh x at 0 1b x b L or 0 0 1 1dd b x b L d where 0M is the slope of the first surface, S0. For 1 1b L x b or surface S1 while the rear axle ( 2C ) is still on the surface S0. Therefore the slope at 2C is between 0M and 1M during this period. According to the geometric relationship in Fig. 6 (a), the slope function is derived as, 1 1tan ( ) sin ( )s M M d x h x L (1) where sx can be solved from 2 2 2 2 1 1 1(1 tan ( )) 2( ) ( ) 0M s d s dx b x x b x L . Then when 2C moves on the surface S2, the slope function is 1( )M d Mh x at 1 2b x b L or equivalently 2 1 2 1 1 3 1cos( )d M b bd b x b L d . The slope function for 2 2b L x b or 2 1 2 1 3 1 1 4 1 1cos( ) cos( )d M M b b b bd b L x b d is obtained from the geometric diagram, Fig. 6 (b), as, 1 2 1 2 1 1 1 1 1 ( ) tan( ) ( ) sin ( )sin( ) cos( ) sin M M d d M M b xh x L b b x b L . (2) Likewise, for the following sections, the slope functions are extracted as, 4 5 2 2 1 2 1 1 3 2 1 1 1 { ( ) 0 - + } cos( ) cos( ) M d M d M M d d h x b b b bb x b b b L , (3) 6 7 3 4 32 1 2 1 3 2 1 3 2 1 1 3 1 { ( ) - + - + } cos( ) cos cos( ) M d M d M M M d d h x b bb b b bb b b x b b b L , (5) 7 8 4 3 2 1 1 3 2 3 3 11 4 3 2 1 3 2 1 3 1 4 3 2 1 3 2 1 3 1 ( )sin cos cos( ) { ( ) sin - + cos cos( ) - + } cos cos( ) d M M M M d d M M d M M d b b b bx b b b h x L b b b bb b b L x b b b bb b b , (6) 8 4 3 2 1 4 3 2 1 3 1 { ( ) 0 - + } cos cos( )M d M d M M d b b b bh x x b b b " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000714_robio.2009.4913083-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000714_robio.2009.4913083-Figure3-1.png", + "caption": "Fig. 3 Geometrical model for supporting anteversion posture", + "texts": [ + " When the McKibben artificial muscle is constricted, inner black circle is forced to rotate with respect to back frame. Waist frame directly connected to the inner black circle is applied and by using thigh belt through connecting belt, waist frame is fixed to the body in order to avoid rotation. In this mechanism, the front part of thigh is burdened by thigh belt during anteversion posture support. Let\u2019s here investigate how much load the front part of thigh has. Geometrical model is shown in Fig.3. O describes the center of rotation at waist. AT [Nm] denotes torque by the upper body as shown in eq.(1), sinMgTA (1) where M [kg] and [m] indicate mass of the upper body and distance from the center of rotation to the gravity center of the upper body. BT [Nm] depicts torque opposes AT as described in eq.(2), tB TbT cos . (2) If AB TT , waist belt stays without revolving. Here let\u2019s calculate T when AB TT , i.e., t t b Mg T TbMg cos sin cossin (3) The load for horizontal and vertical direction of the front part of thigh ( xT and yT respectively) are as follows; sinsin cos sin sin Mg cb Mg TT t t tx (4) sincos cos sin cos Mg bb Mg TT t t ty (5) As a result, one leg receives 2xT and 2yT " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure14-1.png", + "caption": "FIG. 14: (A) SWEPT ROTARY PROFILES (TOOL PATH FOR THE SQUARE FLANGE SHAPE \u2013 WITH ADDITIONAL STOCK TO BE REMOVED, AND A STOCK MODEL FOR THE CAM SHAPE), (B) THE VIRTUAL REPRESENTATION OF THE MACHINED MODELS FOR THE FLANGE AND CAM.", + "texts": [ + " Ideally, rotary slicing, or for the general solution additive surfacing tool paths, are employed for this problem type, especially for 3D fillet geometry. (a) (b) 7 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use The bullet component\u2019s flange corners could be considered protrusions on a rotary base; consequently, a rotary flange could be created, and sets of protrusions built up. However, to avoid issues with segmentation, an hybrid manufacturing approach can be taken. For the non-symmetric rotational components shown in Fig. 14, simple fill rotational additive tool paths and stock models are created based on the swept profile of the geometry (Fig. 14 (a)). Then machining tool paths are created using conventional CAM tools, and the results simulated, as shown in Fig. 14 (b). As with all fabrication process plans, the machine selection, set up strategy, the operation lists, the operation parameters and tool paths, material handling, and so forth need to be determined. Interlacing machining with planar and rotary AM operations increases the complexity of the planning solution space. CAM software to support rotary motion and deposition tool paths is developed. The general process flow for rotary AM process planning is illustrated in Fig. 15. In concept, the rotary positioning is aligned with planar deposition solutions, and the rotary related cases are distinct, such as the decisions with respect to the mandrel / build cylinder, the creation of geometry to support roughing and finishing operations, and build strategies to address nonsymmetric rotary components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001134_2008-01-2623-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001134_2008-01-2623-Figure2-1.png", + "caption": "Figure 2: Bracket to be investigated", + "texts": [ + " It is concluded that ignoring one of those sources of vibration, may result in misinterpretation of fatigue durability of the part. In this study, the fatigue durability of a bracket, which is connected to the chassis and the heater rail system in a inline 4 cylinders diesel engine, is investigated. Engine speed range is 900 rpm \u2013 4500 rpm. 2nd, 4th, and 6th order engine loads are considered to be critical. Finite element model of the whole system is seen in Figure 1, whereas the bracket of interest can be seen in Figure 2. Heater Rail system is connected to the chassis at point 1 and to the engine at point 2. 275 SAE Int. J. Commer. Veh. | Volume 1 | Issue 1 FE MODEL DETAILS - Pipes are modeled with firstorder quadrilateral shell elements. Brackets are modeled with either first-order hexahedral solid elements or firstorder quadrilateral shell elements. Pipe-to-pipe connections are made with spider rigid elements and pipe-to-bracket connections are modeled with rigid elements with single dependent nodes. Total number of nodes in the model: 84584 Total number of elements in the model: 80090 Number of nodes in the bracket (output set): 2344 Number of elements in the bracket (output set): 1545 A modal analysis is performed first in order to obtain the critical modes of the bracket" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002954_042013-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002954_042013-Figure3-1.png", + "caption": "Figure 3. Change of the laser beam relative to the lens position,", + "texts": [ + " The main components were: - a FANUC industrial articulated robot; - an LS-3 solid-state fiber laser by IRE-Polyus ( Fryazino, Moscow Oblast), a subsidiary of IPG Photonics, with the maximum power of 3 kW 1070 nm longwave; - a cladding head with triple-nozzle powder feed to the melt pool by Fraunhofer; - two-hopper powder feeder by Plakart. It should be noted that the main end effector that ensures the process of direct gas-powder laser deposition is the cladding head. It includes a manually controlled collimator for dynamical adjustment of laser beam divergence during operation. Changing the beam divergence will change the focal position of the objective lens (Fig. 3). IPDME2018 IOP Conf. Series: Earth and Environmental Science 194 (2018) 042013 IOP Publishing doi:10.1088/1755-1315/194/4/042013 where 1 is the the top position: divergent beam; 2 is the mid This way, there appears the possibility to measure the diameter of t local melting and the volume of the micro The second stage consisted of the practical development of the modes of direct gas deposition with different diameters of the laser spot. The Inconel 625 initial powder material. Its chemical composition as declared by the manufacturer and the grain size composition identified by laser diffraction using the Analysette 22 particle size measuring unit are presented in Table 1 and Table 2 respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001751_j.jsv.2018.04.026-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001751_j.jsv.2018.04.026-Figure1-1.png", + "caption": "Fig. 1. Configurations of projectile-to-band impact system and propagations of additional incident and reflected waves at a certain instant for (a) central impact and (b) non-central impact.", + "texts": [ + " (ii) A non-central impact inevitably results in asymmetric wave propagations, therefore the projectile would be subjected to a combination of translational and rotational motion. From analytical solutions, we investigated the influence of parameters on the swing amplitude of the projectile. Based on FEM, several efficient and reasonable simulation programs were elaborately developed using LS-DYNA codes, and the simulation results qualitatively or quantitatively showed a satisfactory agreement with our 1-D theoretical results. A projectile-to-band impact system is considered in Fig. 1(a). A projectile, being of a mass Mp, performs a transverse (normal) impact onto the band with an initial velocity V, the front (X-directional) width of the contact zone is D. As the initial condition, the string-like band is fixed by the two supports and is imposed upon a longitudinal pulling force P. It has a length L\u00fe D (the distance between two supports) and amass density per unit length r. The motion of the contact zone of the band is not discussed as it has no influence on the response of the projectile", + " For a random event of normal impact, the impacted location of the band is more likely to deviate from the mid span; therefore, we perform the relevant investigation under non-central impact. Naturally, we can visualize that a complex separation process is featured by the separation of a single bottom edge or by the sequential separations of two bottom edges from the band, and it is also complex to establish resumed contact. However, separation and re-contact are not of concern in the following investigation. The corresponding configuration is shown in Fig. 1(b). The unequal lengths of the two segments isolated by the projectile inevitably result in the asymmetric wave propagations and the two unequal transverse forces at x \u00bc 0. Therefore, the eccentric distance E and the non-dimensionalised eccentric ratio e \u00bc E=~L \u00bc 2E=L \u00f0 1< e<1\u00de have to be introduced, and they are assumed to be positive for a rightward deviation. The dimensional transverse displacement is W\u00f0X; T\u00de, then the dimensionless representation isw\u00f0x; t\u00de \u00bc W= ~W with ~W \u00bc V~T . The front width of the contact zone is D and d \u00bc D= ~W " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000124_pime_auto_1964_179_013_02-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000124_pime_auto_1964_179_013_02-Figure3-1.png", + "caption": "Fig. 3. Diagram of damper velocities. The diagram shows the shock absorbers of the beam axle suspension separated from Fig. 1 for clarity", + "texts": [ + " of the body is not coincident with the centre of rotation 0,, body roll will lower the c.g. of the body in the example shown. This causes a reduction in potential energy of MgH( 1 - cos 0) where H is the vertical height of the c.g. above O1. P = +KJZ- + is(e - .)]a + :K,[z- x - iSp- 4 1 2 +3K,(Y--)2+3KI(aRO+W-A1)2 +tk,(a& +w- A,)' + + h ( x + aL-f if)' +~K,(x-a(yL--f2i)2--gH(i -cOs e) Dissipative function In many vehicles the suspension dampers are not mounted vertically but inclined at an angle 4, Fig. 3. The relative displacement of the two halves of the damper may be obtained by considering the line of action of the damper to be a perpendicular distance (R,) from the axis of body movements (01), and another distance (R2) from the axis of axle movements (02). Resolving the possible displacements of the body and axle along the line of action, the displacement of the left hand damper is: 6, = (RlO- Y cos ++Z sin +)-(R2a-w cos C + x sin 4) differentiating with respect to time: 8, = (&e+R,d- ir cos ++ Y$ sin ++Z sin ++z$ cos 4) Vol179 Pt 2A No 3 at UNIV OF CINCINNATI on June 4, 2016pad" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001483_978-3-319-72730-1_17-Figure17.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001483_978-3-319-72730-1_17-Figure17.5-1.png", + "caption": "Fig. 17.5 bp(x, t) and a(x, t) for leading current", + "texts": [ + " In addition, except for permanent magnet excitation, there is a flux control circuit (not shown), with the desired rotor excitation (rotor flux level \u03a6w p ) depending on the speed (rated rotor flux for speeds below the rated speed and field weakening for higher speeds). 2In the absence of saturation. The desired current magnitude depends on the desired torque: for example, the latter is determined by an external speed control circuit (not shown). The desired displacement between field and current layer is zero for field orientation. However, for a CSI-supply with load commutation, the current should lead the voltage and therefore the current layer should lead the field (Fig. 17.5). In that case, there is vector control and no ideal field orientation. It is obvious that for a supply by a CSI with forced commutation, field orientation (with \u03d1 = 0) is the preferred control. Moreover, for a VSI supply, there is no requirement for a leading current. The control principle illustrated in Fig. 17.6 uses a VSI with a current control loop, Here, the separate control loops formagnitude and angle of the current are replaced by a (real-time) control with the desired current amplitude and current angle as input and the desired switching angles of the PWM-VSI as output (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001035_aqtr.2008.4588797-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001035_aqtr.2008.4588797-Figure1-1.png", + "caption": "Figure 1. Bottom view of the mover", + "texts": [ + " INTRODUCTION Magnetic levitation appears to be a promising technology for achieving the extreme precision required by today\u2019s fabrication processes, especially those used in IC production. Unlike the conventional systems that use stacked movers to obtain the desired DoFs, in case of a magnetic levitated stage, motion on three linear axes along with three rotational ones is obtained with a single moving part, floating over an active surface. From the point of view of motion generation, the mover can be passive, with no need for energy to be brought on it. This is obtained by attaching a permanent magnet (PM) array on its bottom (Figure 1). The pseudo-sinusoidal magnetic field yielded by the PM array interacts with that of the coils, which are mounted in a chessboard array in the fixed part of the machine (Figure 2). All the coils contribute to levitation and motion in the three vertical modes (Z, \u0398x, \u0398y). Half of the coils provide motion in X direction, the other half in Y direction and together, allow the control of yaw (\u0398z). motor pitch) The experimental setup is shown in Figure 3. It is easy to notice that, in this system, the travel range of the mover is restricted only by the size of the coil array, which, based on the modularization and power management methods presented in this paper, is theoretically unlimited", + " When vacuum operation is not a condition, in order to reduce current in coils, it\u2019s worth considering adding air bearing levitation to magnetic levitation. With the appropriate metrology, positioning accuracy of few nanometers can be achieved. In the current embodiment, laser interferometers are used for the horizontal modes and capacitive gap sensors for the vertical ones. Since the plate which cover the coils is not a reliable target for the gap sensors mounted on the bottom of the mover (see Figure 1), we plan to find another solution for sensing the vertical modes. The great and variable number of coils leads to the necessity of a modular control and driving system. This is more obvious when a structure like that shown in Fig. 4 is considered. Those presented in the following sections refer to the embodiment shown in Fig. 3, but some ideas can be easily ported to the structure shown in Fig. 4. The control strategy is presented in Section II. The modular control and driving structure is presented in Section III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002805_detc2018-85407-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002805_detc2018-85407-Figure10-1.png", + "caption": "Figure 10. The space location of kingpin", + "texts": [], + "surrounding_texts": [ + "In above equations, \ud835\udeff\ud835\udc611_\ud835\udc5f represents the steering front\nturning angle at this time.\nThe initial process of the wheel positive rebound\nThe upper part of tire rebound because of the elastic force. But the elastic force is less than the maximum static friction. So the tire part contacting with the ground doesn\u2019t turn. Equation (8) and (9) describe the dynamical model. And the figure (7) shows the diagram of the model.\npositive rebound\n\ud835\udc3d1?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc61?\u0307?\ud835\udc61 + \ud835\udc3e\ud835\udc61\ud835\udeff\ud835\udc61 + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 (8)\n(\ud835\udc35\ud835\udc61?\u0307?\ud835\udc61 + \ud835\udc3e\ud835\udc61\ud835\udeff\ud835\udc61) \u2264 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 (9)\nEquation (10) and (11) describe the dynamical model of the\nfinal time of this process.\n0 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 (10)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 = \ud835\udc58\ud835\udc61 \u2219 (\ud835\udeff\ud835\udc611_\ud835\udc53 \u2212 \ud835\udeff\ud835\udc500) (11)\nIn above equations, \ud835\udeff\ud835\udc500 represents the turning angle of the tire part contacting with the ground at this time, \ud835\udeff\ud835\udc611_\ud835\udc53 represents the front wheel turning angle at this time.\nThe middle process of the wheel positive rebound\nEquation (12) describes the dynamical model. And the figure\n(8) shows the diagram of the model.\npositive rebound\n{ \ud835\udc3d1 \u2219 ?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc61 \u2219 (?\u0307?\ud835\udc61 \u2212 ?\u0307?\ud835\udc50) + \ud835\udc3e\ud835\udc61(\ud835\udeff\ud835\udc61 \u2212 \ud835\udeff\ud835\udc50) + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60\n\ud835\udc3d\ud835\udc612 \u2219 ?\u0308?\ud835\udc50 + \ud835\udc35\ud835\udc61 \u2219 (?\u0307?\ud835\udc50 \u2212 ?\u0307?\ud835\udc61) + \ud835\udc3e\ud835\udc61(\ud835\udeff\ud835\udc50 \u2212 \ud835\udeff\ud835\udc61) = \u2212\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50\n(12)\nIn above equation, \ud835\udc3d\ud835\udc612 represents the moment of inertia of the tire part contacting with the ground, \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 represents the friction torque produced by the sliding friction force between road and tire.\nThe final process of the wheel positive rebound\nEquation (13) and (14) describe the dynamical model of the final time of this process. And the figure (9) shows the diagram of the model.\npositive rebound\n0 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 (13)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 = \ud835\udc58\ud835\udc61 \u2219 (\ud835\udeff\ud835\udc612_\ud835\udc53 \u2212 \ud835\udeff\ud835\udc501) (14)\nIn above equations, \ud835\udeff\ud835\udc501 represents the turning angle of the tire part contacting with the ground at this time, \ud835\udeff\ud835\udc612_\ud835\udc53 represents the front wheel turning angle at this time.\n4 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "When vehicle is stationary, the self-aligning torque is mainly caused by vertical force. Because of the existence of the kingpin caster angle and kingpin inclination angle, the vertical force acting on the tire contact area produces self-aligning torque around the kingpin. Figure (10) shows the space location of kingpin. When the front wheel turns, the relative position between kingpin and vertical force changes and the selfaligning torque changes. The literature [10] uses the homogeneous coordinate transformation to calculate the selfaligning torque. And the result shows that the total self-aligning torque produced by the vertical force is approximately proportional to the front wheel steering angle. And Eq. (15) is used to calculate the self-aligning torque in this paper.\n\ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b = \ud835\udc58\ud835\udc4e \u2219 \ud835\udc39\ud835\udc67 \u2219 \ud835\udeff\ud835\udc61 (15)\n\ud835\udf37 is kingpin inclination angle, \ud835\udf38 is kingpin caster angle, \ud835\udc6d\ud835\udc9b is\nthe vertical force\nWhen vehicle steers, the friction occurs between tire and road surface. The contact patch is approximate a circle if the tire bears light road. With the load increasing, the entire width of the tire contacts with the road surface, and contact patch becomes rectangle at this time. The shape of contact patch is different when the tire pressure is different. The contact patch is oval at normal tire pressure. The empirical formula (16) is used in this paper [11].\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 = \ud835\udc58\ud835\udc53 \u2219 \ud835\udf07 \u2219 \ud835\udc39\ud835\udc67\n1.5\n\u221a\ud835\udc5d (16)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 \u2264 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 (17)\nIn above equations, \ud835\udc58\ud835\udc53 is a constant which equates to 0.013, \ud835\udf07 is the friction coefficient which equates to the road adhesion coefficient.\nThe model of viscous friction torque can be simplified as\nequation (18). And the \ud835\udc58\ud835\udc59 is a constant in the equation.\n\ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 = \ud835\udc58\ud835\udc59 \u2219 \ud835\udc60\ud835\udc54\ud835\udc5b(\ud835\udeff\ud835\udc61) (18)\nThe middle process of the wheel positive rebound is used to estimate the road adhesion coefficient. The moment of inertia of the tire part contacting with the ground is small that it can be ignored in equation (12). And the dynamical model can be rewrite as equation (19).\n\ud835\udc3d1 \u2219 ?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 (19)\nTaking the equation (15), (16) and (17) into equation (19).\n\ud835\udc3d1 \u2219 ?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \ud835\udc58\ud835\udc4e \u2219 \ud835\udc39\ud835\udc67 \u2219 \ud835\udeff\ud835\udc61 \u2212 \ud835\udc58\ud835\udc59 \u2219 \ud835\udc60\ud835\udc54\ud835\udc5b(\ud835\udeff\ud835\udc61) \u2212 \ud835\udc58\ud835\udc53 \u2219 \ud835\udf07 \u2219 \ud835\udc39\ud835\udc67\n1.5\n\u221a\ud835\udc5d\n(20)\nIn equation (20), \ud835\udc3d1and \ud835\udc35\ud835\udc60 are characteristic parameters of steer system which can be measured. The adhesion coefficient \ud835\udf07 can be obtained through estimating \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 .\nEstablishing the Luenberger reduced-order disturbance\nobserver\nSelecting the state variables:\n{\n\ud835\udc651 = \ud835\udeff\ud835\udc61 \ud835\udc652 = ?\u0307?\ud835\udc61\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50\n(21)\nAnd the Eq. (19) can be rewritten as:\n?\u0307?2 + \ud835\udc35\ud835\udc60\n\ud835\udc3d1 \u2219 \ud835\udc652 \u2212\n\ud835\udc58\ud835\udc4e\u2219\ud835\udc39\ud835\udc67\n\ud835\udc3d1 \u2219 \ud835\udc651 = \u2212\n\ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60\n\ud835\udc3d1 \u2212\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50\n\ud835\udc3d1 (22)\nThe state equation can be written as:\n5 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "[\n?\u0307?1 ?\u0307?2\n?\u0307?\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50\n] = [\n0 1 0\n\u2212 \ud835\udc58\ud835\udc4e\u2219\ud835\udc39\ud835\udc67\n\ud835\udc3d1\n\ud835\udc35\ud835\udc60 \ud835\udc3d1 \u2212 1 \ud835\udc3d1\n0 0 0\n] [\n\ud835\udc651 \ud835\udc652\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 ] + [ 0 1 0 ] [\u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 \ud835\udc3d1 ] (23)\n\ud835\udc66 = [ 1 0 0 0 1 0 ] [\n\ud835\udc651 \ud835\udc652\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50\n] (24)\nThe state equation is discretized as:\n[ \ud835\udc651(\ud835\udc58 + 1) \ud835\udc652(\ud835\udc58 + 1)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58 + 1) ] = [\n1 \ud835\udc47\ud835\udc60 0\n\u2212 \ud835\udc47\ud835\udc60\u2219\ud835\udc58\ud835\udc4e\u2219\ud835\udc39\ud835\udc67\n\ud835\udc3d1 (1 +\n\ud835\udc47\ud835\udc60\u2219\ud835\udc35\ud835\udc60\n\ud835\udc3d1 ) \u2212\n\ud835\udc47\ud835\udc60 \ud835\udc3d1\n0 0 1\n] [ \ud835\udc651(\ud835\udc58) \ud835\udc652(\ud835\udc58)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) ] +\n[ 0 \ud835\udc47\ud835\udc60\n0 ] [\u2212\n\ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60\n\ud835\udc3d1 ]\n(25)\n\ud835\udc66(\ud835\udc58) = [ 1 0 0 0 1 0 ] [\n\ud835\udc651(\ud835\udc58) \ud835\udc652(\ud835\udc58)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) ] (26)\nIn the equation, \ud835\udc47\ud835\udc60 is the sampling period. Equation (26) can be written as:\n[ \ud835\udefe1(\ud835\udc58 + 1)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58 + 1) ] = [\n\ud835\udc3411 \ud835\udc3412 \ud835\udc3421 \ud835\udc3422 ] [ \ud835\udefe1(\ud835\udc58) \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) ] + [ \ud835\udc351 \ud835\udc352 ] \ud835\udc62 (27)\n\ud835\udc66(\ud835\udc58) = [\ud835\udc361 \ud835\udc362] [ \ud835\udefe1(\ud835\udc58)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) ] (28)\nThe reference model can be written as:\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58 + 1) = \ud835\udc3422 \u2219 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) + \ud835\udc3421 \u2219 \ud835\udc66(\ud835\udc58) + \ud835\udc352 \u2219 \ud835\udc62 (29)\n\ud835\udc67 = \ud835\udc3412 \u2219 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) (30)\nThe observer is established as\n?\u0302?\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58 + 1) = (\ud835\udc3422 \u2212 \ud835\udc3f \u2219 \ud835\udc3412) \u2219 ?\u0302?\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) + \ud835\udc3421 \u2219 \ud835\udc66(\ud835\udc58) + \ud835\udc352 \u2219\n\ud835\udc62 + \ud835\udc3f \u2219 \ud835\udc66(\ud835\udc58 + 1) \u2212 \ud835\udc3f \u2219 \ud835\udc3411 \u2219 \ud835\udc66(\ud835\udc58) \u2212 \ud835\udc3f \u2219 \ud835\udc351 \u2219 \ud835\udc62\n(31)\nIn order to take place of \ud835\udc66(\ud835\udc58 + 1) which can\u2019t be detected,\n\ud835\udc64(\ud835\udc58) is introduced.\n\ud835\udc64(\ud835\udc58) = ?\u0302?\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58) \u2212 \ud835\udc3f \u2219 \ud835\udc66(\ud835\udc58) (32)\nAnd\n\ud835\udc64(\ud835\udc58 + 1) = ?\u0302?\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50(\ud835\udc58 + 1) \u2212 \ud835\udc3f \u2219 \ud835\udc66(\ud835\udc58 + 1) (33)\nIn equation (33):\n\ud835\udc350 = \ud835\udc352 \u2212 \ud835\udc3f \u2219 \ud835\udc351 (35)\n\ud835\udc38 = (\ud835\udc3422 \u2212 \ud835\udc3f \u2219 \ud835\udc3412) \u2219 \ud835\udc3f + \ud835\udc3421 \u2212 \ud835\udc3f \u2219 \ud835\udc3411 (36)\nAnd the \ud835\udc3f in above equation is error feedback gain matrix. (\ud835\udc3422 \u2212 \ud835\udc3f \u2219 \ud835\udc3412) is asymptotically stable matrix. The eigenvalue of (\ud835\udc3422 \u2212 \ud835\udc3f \u2219 \ud835\udc3412) which can be changed by adjusting \ud835\udc3f influences the rate of convergence. And the constraint condition is\n|\ud835\udc3422 \u2212 \ud835\udc3f \u2219 \ud835\udc3412| < 0 (37)\nMETHOD VALIDATION\nThe verification test is conducted in the vehicle equipped with a SBW system which has been described in the introduction. The test conditions are ice surface and snow surface road. Table 1 shows the parameters value of the experiment. The sampling period is 10ms. During the test, the front wheel is turned to three target angles, and then steer torque is released suddenly. The convergent value of estimating road adhesion coefficient is showed in table 2.\nTable 1. The parameters of the experiment\nVertical Force \ud835\udc6d\ud835\udc9b \ud835\udfd1\ud835\udfd6\ud835\udfd0\ud835\udfd3 \ud835\udc0d\nThe equivalent moment of inertia of\nthe steer system \ud835\udc71\ud835\udfcf\n0.02 kg/m2\nThe coefficient of viscous friction\ntorque \ud835\udc8c\ud835\udc8d\n0.7 Nm\nThe coefficient of self-aligning\ntorque \ud835\udc8c\ud835\udc82\n4.65 \u00d7 10\u22123\nThe damping coefficient of the supporting mechanism in steer\nsystem \ud835\udc69\ud835\udc94\n2.9 Nms/rad\nTire Pressure \ud835\udc77 230 kPa\nTable 2. The results of the experiment\nRoad\nSurface\nEstimated\nvalue of\nEstimated\nvalue of\nroad\nActual value\nof road\n6 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use" + ] + }, + { + "image_filename": "designv11_92_0000881_pi-b-2.1962.0229-Figure19-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000881_pi-b-2.1962.0229-Figure19-1.png", + "caption": "Fig. 19.\u2014Localizer and v.o.r. aerial for use on narrow fins.", + "texts": [ + " Since the height of the system above the fuselage may be considerable, a much higher proportion of radiation occurs in the azimuthal plane with a consequent increase in range. Fig. 18 shows azimuth patterns for a fin-mounted and a pedestal-mounted aerial, plotted to the same scale. If the fin is narrow, say not more than 6 in wide at the point of mounting, it is possible to use a V-dipole mounted inside the fin with the arms projecting through insulating covers in the sides of the fin. The dipole can be fed from a conventional balance unit bolted to the inboard ends of the elements, as shown in Fig. 19. On wider fins the aerial arms are separate Pedestal. Fin. unipoles which may be bent-sleeve aerials as used for v.h.f. communication. Two methods of feeding are then possible: a pair of coaxial cables of equal electrical length connected to a balance unit, or two cables differing by a half-wavelength and connected to a T-junction. Since the difference will be exactly a half-wavelength only at the mid-band frequency, the latter method can be used only if the aerials, individually, have a low v.s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002082_1.4040813-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002082_1.4040813-Figure12-1.png", + "caption": "Fig. 12 Pressure profiles versus operating speed, 124 kPa specific load, 100% flow rate", + "texts": [ + " The higher damping ratios of the starved model can also be attributed to this decrease in bearing stiffness. Overpredicted bearing damping in both models might also explain the higher damping ratios of the starved model. At generally higher damping values, changes in stiffness are more influential in the resulting damping ratio, as observed in the two models. Generally, lower bearing damping values in both models would act to bring the damping ratio predictions of the starved model closer to those of the flooded model in the presence of lower bearing stiffness. Figure 12 illustrates the predicted pressure profiles of the starved bearing model over the higher operating speed range for the nominal load and flow conditions. Leading edge cavitation can be seen in the upper pads that increases in severity with increasing speed, lessening the load on the lower pads. This progressive decrease in the scope and magnitude of the developed hydrodynamic film layer results in the decrease in bearing stiffness and damping seen in Fig. 11. The experimentally identified damped natural frequencies agree very well with the starved model predictions, while the identified damping ratios are consistently lower than the model predictions", + " For all oil flow conditions, the starved model predicts lower stiffness and damping values than the flooded model across all speeds. Higher predicted damping ratios in the starved model suggest that the lower stiffness values are, once again, defining the trends seen in the resulting modal properties. Figure 18 illustrates the predicted pressure profiles of the starved bearing model for varying oil supply flow rates under the nominal load condition at 10,000 rpm. Similar to the speeddependent pressure profiles in Fig. 12, cavitation can be seen in the upper pads that increases in severity with decreasing oil supply flow rate, lessening the load on the lower pads. Again, a decrease in hydrodynamic film development in the upper pads and reduced loading of the lower pads result in the decreased bearing stiffness and damping seen in Figs. 16 and 17. As with the loaddependent observations, the models and data produce opposite trends. While the data clearly show a decrease in damping ratio, or system stability, with decreasing flow rate, both models predict an increase in stability with decreasing flow rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002374_icca.2018.8444305-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002374_icca.2018.8444305-Figure1-1.png", + "caption": "Fig. 1. The original sector region of sensor si is mapped into a circle.", + "texts": [ + " Let pi = [xi yi \u03b8i]T \u2208 R3 denote the ith sensors center coordinate and the orientation angle in radian, and Ps = {p1, p2, \u00b7\u00b7\u00b7, pn}, with n being the total number of sensors be the set of all the sensors. dE(,) stands for the Euclidean distance on R2. The considered 2D sensors have a circular sector sensing region, denoted by Si \u2282 R2 for sensor i, with the boundary \u2202Si. Let r > 0 be the radius of the sector, \u03b1 the central angle of the circular sector in radians. In this subsection, we aim to construct a homeomorphism Hi: R2 \u2192 R2 to map the sector region Si into a unit circle. Without loss of generality, as shown in Figure 1(a), we choose the middle point oi of the central axis within the sector as the \u2018sensing center\u2019. F(XoY) denotes the world coordinate system, Fi(XioiYi) is the ith sensor frame with its origin located at oi, and its Yi axis is aligned with the central axis of the sector. Fti(XtiotiYti) denotes the frame in the transformed space through the homeomorphism Hi, as shown in Figure 1(b) and (c) respectively. 1) Relationship between F and Fi: in the frame F, the position of the sensor is represented by ci = [xi yi]T \u2208 R2, while its orientation is \u03b8i. ci and \u03b8i constitute the configuration of the ith sensor as pi = [cTi \u03b8i]T = [xi yi \u03b8i]T . The relationship between F and Fi is expressed as qs = RT (q \u2212 T ), (1) where q and qs are the coordinates of an arbitrary point expressed in F and Fi, respectively, and R = [ cos \u03b8i \u2212 sin \u03b8i sin \u03b8i cos \u03b8i ] , T = [ xi + r 2 cos \u03b8i yi + r 2 cos \u03b8i ] " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure8-1.png", + "caption": "Fig. 8 Comparison of the results of displacement during acceleration. (a) Using the spring element-based method; (b) using the plate theory-based method", + "texts": [ + " Lastly, explaining the simulation results using the spring element method is difficult, whereas using the newly proposed method solves the usual problems related to the DOF coupling effect. Figure 7 shows the results of using the two methods to calculate the rotor-bearing system\u2019s displacement and deformation during acceleration. A comparison of the results obtained by Methods 1 and 2 is shown in Figs. 8(a) and 8(b), respectively. When the spring element-based method is used, the maximum displacement is in the circumferential direction (Fig. 8(a)), which is obviously not consistent with reality. However, when the proposed plate theory-based method is used under the same load conditions, the direction of maximum displacement (0.051 mm) is in the radial direction (direction of resultant acceleration), which is consistent with reality (Fig. 8(b)). 4.2 Effect of displacement parameters To improve the application of the proposed method, we can develop a database or software tool for the accurate selection of displacement parameters. For an annular plate, the outer radius is limited by the radius of the bearing hole, inner radius b is limited by the radius of the rigid shaft, and Poisson\u2019s ratio \u03bc is limited by the plate material. Hence, only the thickness of the annular plate needs to be changed to determine the correct displacement parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003041_apeie.2018.8545424-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003041_apeie.2018.8545424-Figure1-1.png", + "caption": "Fig. 1. The piecewise homogeneous zone named \u201cmagnets-poles\u201d.", + "texts": [ + " This method allows transform the analytic solution of the partial differential equation into four-terminal circuit equations and synthesize the cascade equivalent circuit of the electric machine. In this paper, a synthesis of an active link of a cascade AH-equivalent circuit of a salient-pole synchronous electric machine is considered. In addition, it should be noted that the exciting field is created by the inductor with tangential permanent magnets made of rare-earth metals. II. AN INDUCTOR OF A SYNCHRONOUS ELECTRIC MACHINE WITH TANGENTIAL PERMANENT MAGNET EXCITATION AND A STANDARD ACTIVE LINK OF A CASCADE A-H-EQUIVALENT CIRCUIT The Fig. 1 shows the piecewise homogeneous zone named \u201cmagnets-poles\u201d where prismatic magnets are purposively configured as sphenoid for lines \u201cmagnet-pole\u201d must coincide with coordinate surfaces of the cylindrical coordinate system. Because of symmetry, let consider the problem solution upon the half of the pole pitch i.e. with the independent coordinate \u03b1 varying from 0 to \u0398. Poles and magnets are attached to a rotor and separated from a shaft with a non-magnetic tube which magnetic reluctance is too high", + " (6) When the rotor of the electric machine is rotating with the angular frequency \u03c9, a detached observer sees traveling waves of the functions (5) and (6) which first harmonics in the complex plane representation are given by the expressions: 1 2 n nA j C r C r Q\u2212 = + , (7) 1 1 1 2 n n i nH j C r C r\u2212 \u2212 \u2212 = \u2212 \u2212 \u03a8 \u03bc , (8) where ( ) 0 0 2 cos cos sin cos i Q n p d K n p d \u03b2 \u0398 \u03b2 \u03bc = \u03b1 \u03b1 \u03b1 + \u0398 \u03bc + \u0398 \u2212 \u03b1 \u03b1 \u03b1 , ( ) 0 0 2 cos cos sin cos i i f n p d K n p d \u03b2 \u0398 \u03b2 \u03bc \u03a8 = \u03b1 \u03b1 \u03b1 + \u0398 \u03bc \u03bc + \u0398 \u2212 \u03b1 \u03b1 \u03b1 \u03bc , 2p is the number of poles. At the boundaries of the piecewise homogeneous zone named \u201cmagnets-poles\u201d where 1r r= and 2r r= (see. Fig. 1), current sheets are located. These current sheets are equal to the coercitivity of the magnets and given by the piecewise continuous function: ( )0 0 0 cH H < \u03b1 < \u03b2 \u03b1 = \u03b2 < \u03b1 < \u0398 . (9) The first harmonic of the function (9) in the complex plane representation is given by the expression: 0 2 sincHH j p p = \u03b2 \u0398 . (10) Let the vector potential and the tangential component of the magnetic intensity vector are known at the boundaries of the piecewise homogeneous zone named \u201cmagnets-poles\u201d (when 1r r= and 2r r= )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002883_aaf0d1-Figure18-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002883_aaf0d1-Figure18-1.png", + "caption": "Figure 18. Predicted bandgap opening in highly-oriented wrinkles formed in chemical vapour deposition grown graphene , Omar M. Dawood et al.", + "texts": [], + "surrounding_texts": [ + "O. M. Dawood is grateful for the support of the Ministry of Higher Education and Scientific Research (MOHESR), Iraq" + ] + }, + { + "image_filename": "designv11_92_0000501_fie.2009.5350615-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000501_fie.2009.5350615-Figure7-1.png", + "caption": "FIGURE 7 EVICE FOR USE IN ), (COURTESY, LE", + "texts": [], + "surrounding_texts": [ + "PORTAB PO Oc ation Confere BLE ULTRASOUND (HIMALAYAS E ULTRASOUND D (DARFUR, SUDAN LE ULTRASOUND OR REGIONS (SIER tober 18 - 21, nce 2009, San An D FOR HIGH ELEVA SEN, SPOFFORD) A TRIAGE TENT DU CLAIRE, MCCORM ANCY AND PEDIA RTESY WILSON, B tonio, TX TION USE RING ACK) TRIC USE IN LAIR) 978-1-4244-4714-5/09/$25.00 \u00a92009 IEEE October 18 - 21, 2009, San Antonio, TX 39th ASEE/IEEE Frontiers in Education Conference M3E-6 This really allowed each team to have a design unique from their classmates, but have to explain how the design addressed the challenges of the user in the situation they imagined. This approach also allowed for students in the class to learn from other teams who took a completely different approach to the project. Another unique learning objective achieved with this project was achieved within the ID-ET teams. The students stated repeatedly how much they learned from their teammate during the project. Engineering Technology students were given experience using the techniques ID students use to develop ideas, such as mind mapping and design for each function. Industrial Design students were given experience in design for manufacturability. Each team had to justify material choices, manufacturing processes, and assembly techniques. One challenge to this approach was that it was not clear which student in each team was responsible for the final deliverables of the project. The PET student was responsible for the CAD drawings included exploded assemblies. The ID student was responsible for putting the CAD designs into a rendered \u201ccontext\u201d that would help to explain the design." + ] + }, + { + "image_filename": "designv11_92_0001483_978-3-319-72730-1_17-Figure17.13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001483_978-3-319-72730-1_17-Figure17.13-1.png", + "caption": "Fig. 17.13 Direct flux control: all sectors and hysteresis band", + "texts": [ + " For the flux vector \u03a8 s in the figure, a voltage vector u3 will cause a counter-clockwise rotation, without a noticeable variation in amplitude. A voltage vector u2 will cause both a counter-clockwise rotation and an increase of the amplitude. When a zero voltage vector is switched, the flux will remain more or less constant (in fact, it will slightly decrease because of the resistance). In the usual implementation of DTC, the complex plane is divided into six sectors of \u03c0/3, each centred around one of the non-zero vectors (see Fig. 17.13). We will now turn to the case where the flux vector is in the first sector (\u2212\u03c0/6 \u2264 \u03d1 \u2264 \u03c0/6) and the required rotation direction is positive (i.e. counter-clockwise), as is the case in (b) in Fig. 17.13. Which vector is switched depends on the deviation of the angle and amplitude of the actual flux vector with respect to the desired one (if hysteresis controllers are used). If the actual flux vector is behind (lags) the desired one, either u2 or u3 can be switched to advance the flux. When the actual amplitude is lower than the desired one, u2 should be switched; when the actual amplitude is higher, u3 is switched. At some point, the actual flux vector may start to lead the desired one. Then, normally one of the zero vectors (u0 or u7) will be switched, causing the actual flux vector to stand still (and slightly decrease). In principle, either one of the zero voltage vectors can be chosen, but the one resulting in the lowest number of switchings will be preferred. To avoid too high a switching frequency, a wide hysteresis band can be utilised (see (a) in Fig. 17.13). Note that for an actual flux vector leading the desired one, u5 or u6 could have been used as well. However, this would increase the switching frequency too much and is thus best avoided. Repeating this for the other sectors and also adding the sequence for the negative rotation direction results in the switching Table17.1. As mentioned above, in steady state, only the first and second rows are used for the positive (counter-clockwise) rotation direction, while only the third and second rows are used for the negative (clockwise) rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000742_6.2009-2404-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000742_6.2009-2404-Figure2-1.png", + "caption": "Figure 2: The UAV Geometry and Mass Distribution", + "texts": [ + " (28) for V\u0303 and arranging the resulting equations, the first-order equations can be cast in the state-space form \u02d9\u0303x(t) = A(t)x\u0303(t) +B(t)u\u0303(t) (31) where x\u0303 = [R\u0303T \u03b8\u0303 T q\u0303T r q\u0303T l \u03be\u0303 T r \u03be\u0303 T l p\u0303T v \u03c1\u0303 T \u03c9 p\u0303T r p\u0303T l \u03c1\u0303 T r \u03c1\u0303 T l ]T (32) is the first-order state, u\u0303 is the first-order control and A(t) and B(t) are the coefficient matrices. A and B are time-invariant if \u03b8\u0302, v\u0302 and \u03c9\u0302 are constant, and are time-varying if otherwise. The numerical model of the UAV includes the geometry, stiffness, mass distribution and aerodynamic data. Aircraft model of Ref. 10 will be used in this paper. Mass distribution is given in terms of distributed mass per unit span and lumped masses. Geometry and mass distribution of the aircraft are given in Fig. 2. Mass centers of the wing and the horizontal and vertical stabilizers are located at the half-chord. Elastic axis of the wing is also located at the half-chord. EIy = 2\u00d7 104 Nm2, EIz = 4\u00d7 106 Nm2, GJ = 1\u00d7 104 Nm2. We assume CL0 = 0 for all lifting surfaces and CL\u03b1i = 2\u03c0ARi/(2 + ARi), i = r, l, h, v. In this numerical example, we assume both right- and left-wing have n bending and torsional degrees of freedom. The damping matrices are assumed to be proportional to the stiffness matrices where the proportionality constants are equal to 2\u03b6 \u221a \u039b in which \u03b6 is a structural damping factor and \u221a \u039b is the lowest natural frequency of the half wing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000968_tmag.2009.2012536-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000968_tmag.2009.2012536-Figure5-1.png", + "caption": "Fig. 5. IPM motor with an overhang rotor.", + "texts": [ + " 4 shows the aspect ratio of the element in the insulation gap and CPU time at the first nonlinear iteration. The aspect ratio is defined as the length ratio of the longest and shortest edges in one element. The element shape is poorer when the value of the aspect ratio is larger. The AMGCG method keeps the better performance for the mesh of poor shaped elements than the standard ICCG method. It is supposed that the AMGCG method is robust to the poor shaped elements. 1) IPM Motor With Overhang Rotor Model: This IPM motor model, shown in Fig. 5, consists of the four interior magnets, 24 slots stator, and the overhang rotor. The stack length of the rotor is 5 mm longer than the stator. The rotor and stator are constructed by laminating steel sheets. The thickness of the steel sheet is 0.5 mm and the space factor is 0.96. The influence of Table III shows the performance of the standard ICCG method and the AMGCG method. The AMGCG method shows better performance than the standard ICCG method. Fig. 6 shows the distribution of magnetic flux density in the Z-axis direction of the rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003942_iccairo.2018.00054-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003942_iccairo.2018.00054-Figure1-1.png", + "caption": "Fig. 1. (Top) Photograph and (bottom) mechanical drawing of MLA.", + "texts": [ + " The intense heat from the Sun requires the spacecraft to point its sunshade toward the Sun at all times; during noon\u2013midnight orbits, this requirement constrains the payload deck to point within \u223c10\u25e6 about the ecliptic south pole, with MLA ranging at a slant angle as high as 70\u25e6. The MLA measures the topography of Mercury via laser pulse time-of-flight and spacecraft orbit position data. The primary science measurement objectives for MLA are to provide a high-precision topographic map of the northern polar regions, to measure the long-wavelength topographic features of the mid-to-low latitude regions, and to detect and quantify the planet\u2019s forced librations. Fig. 1 shows a photograph and a mechanical drawing of MLA. Fig. 2 shows U.S. Government work not protected by U.S. copyright. the MESSENGER payload instruments and the location of MLA on the spacecraft. MLA operates under a harsh and highly dynamic thermal environment, due to the large variation in heat flux from the Mercury surface from daytime to nighttime and from deep space background. The transmitter and receiver optics undergo rapid and uneven swings in temperature during science measurement (tens of degrees per hour at the laser-beam expander and the receiver telescope)", + " The calibration results are described here. The MLA coordinates, with respect to the spacecraft coordinate system, are obtained from a combination of the MLA measurements and spacecraft survey results after payload integration. The definitions of the MESSENGER spacecraft coordinate system are labeled in Fig. 2. The XY plane is the same as that defined by the launch vehicle separation plane, and the origin is at the center of the adapter ring. MLA orientation in the spacecraft coordinate system is shown in Fig. 1. All the MLA physical dimensions and angular directions can be referenced to the alignment reference cube attached to the side of the main housing, as shown in Fig. 1. The coordinates of the MLA alignment reference cube in the MESSENGER spacecraft coordinate system have been measured to be [\u221215.2, 63.5,\u221218.7] mm. The estimated spacecraft center of mass was at [0, 8,\u2212827] mm, with an uncertainty of [\u00b110,\u00b110,\u00b120] mm at launch. The center of mass during the mission is a function of propellant usage and distribution within the fuel tanks. The MLA range measurement is referenced to the XY plane that passes through the center of the reference cube. The MLA laser pointing angle was measured by focusing the beam with an off-axis parabola (OAP) to a reticle and measuring the angular offset between the reticle and the OAP optical axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000918_icma.2009.5246262-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000918_icma.2009.5246262-Figure5-1.png", + "caption": "Fig. 5 Simplified model of robot.", + "texts": [ + " For instance, we can specify the number of robots , the positions of static obstacles, the position of each robot's goal, and the initial positions for each robot et al. The simulation environment is a two dimension space of 900 pixels by 600 pixels. The size of a single robot is 30 pixels, and all robots in the system have the same speed of 1.5 pixels per second. A. The Realization ofTrajectory Tracking For tracking the sub-goal of the robot and counting the position and pose of the robot real-timely, in this paper we adopt the model of two-wheel differential driven which is shown as Fig. 5. R EFERENCES we conducted experiments for different scenarios. Simulation results show two points: the first is that these algorithms are effective and useful, the second is that these algorithms improve the safety and smoothness of path planning. Our future work is to apply the developed algorithms in a real multi-robot system. [I) T. Xu and Z. Tang, \"Research on obstacle avoidance planning for mobile robots in dynamic world,\" Robots, vo!. 25, no. 2, pp. 117-122, March 2003. (2) F. Zhang and D" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000398_s102833580906010x-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000398_s102833580906010x-Figure1-1.png", + "caption": "Fig. 1. Inverted pendulum with the periodic excitation of the base.", + "texts": [ + " a w1 TB0M 1\u2013 B0w1 w1 TC1w1 ------------------------------------\u2126 2\u2013\u2013 O \u2126 4\u2013( ).+= a \u2126 2\u2013 w1 TC1w1 ------------------- w1 TB0wk( )2 k 1= n \u2211\u2013 O \u2126 4\u2013( ),+= q\u0307\u0307 p\u2013 \u03b4b t( )+( )+ q 0,= obtain the stabilization region in the first approximation as (31) For T = 2\u03c0, this formula coincides with that obtained previously in [10]. If b(t) = cost, a = 1, which is wellknown. As the second example, we consider the inverted double pendulum consisting of two concentrated masses m1 and m2 connected by rigid massless rods of identical length l and rigidity c in the gravity field (Fig. 1). The Lagrange function L = K \u2013 V of the system is determined by the expressions (32) where g is the acceleration of gravity. We consider the periodic excitation of the base z = acos\u2126t. Then, according to the d\u2019Alembert principle, it is necessary to substitute g + instead of g. The equations of motion p a\u03b42 2 ------- , a< 4 T -- b t( ) b \u03c4( )\u03c4 \u03c4d td 0 t \u222b 0 T \u222b= \u2013 2 T2 ----- b t( )t td 0 T \u222b\u239d \u23a0 \u239c \u239f \u239b \u239e 2 . K 1 2 -- m1 m2+( )l2\u03b8\u03071 2 1 2 --m2l2\u03b8\u03072 2 += + m2l2\u03b8\u03071\u03b8\u03072 \u03b81 \u03b82\u2013( ),cos V c\u03b81 2 2 ------- c \u03b82 \u03b81\u2013( )2 2 -------------------------+= + m1 m2+( )gl \u03b81 m2gl \u03b82,cos+cos z\u0307\u0307 298 DOKLADY PHYSICS Vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure16-1.png", + "caption": "Fig. 16 i-foot exterior. (a) Front view. (b) Side view", + "texts": [ + " Hence, in the designing process, we considered safety and reliability as the basic performance specifications as long as walking ability. ifoot is designed in an eggshell-like shape that wraps around a driver and that can guarantee a driver\u2019s safety in worst-case scenarios. Although many other mountable robots have been proposed so far [4,13,32,43], i-foot is the only mountable, walking robot with safety and reliability. In addition, we were also particular about its user interface. i-foot is designed so that people could easily mount/dismount and that anyone could easily drive. Figure 16a shows a front view of i-foot and Fig. 16b shows a side view. Table 1 lists the main specifications. i-foot has a single seat with seatbelts. To guarantee the driver\u2019s safety, the cabin is comprised of a bucket-type seat that covers body parts including the arms, hips, and head and is designed so that the driver\u2019s body Table 1 i-foot specifications Height [mm] 2360 Length [mm] 1500 Width [mm] 910 Weight [mm] 200 Joint DOF 12 Weight capacity [kg] 60 Seating capacity 1 Rated walking speed [km/h] 1:35 does not stick out. The cabin is also designed to be able to absorb the impact and protect a driver even if a large impact occurs by any chance, such as falling over or a collision with an obstacle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001617_s40799-018-0232-7-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001617_s40799-018-0232-7-Figure8-1.png", + "caption": "Fig. 8 Location of the attached strain gage on the solid cylinder (mm)", + "texts": [ + "185, as well as transverse sensitivity coefficients of 1.0 and 0.7, respectively; both parameters values are provided by the manufacturer. The attachment of the strain gage to the solid cylinder is based on the traditional methodology of adhesive application [29]. Figure 7 shows the gage properly placed on and connected to the solid cylinder. Note that the location of the gage is the same that is considered in the numerical analysis, on the upper part of the cylinder and at 50 mm from the applied load, as shown by Fig. 8. This location is assumed to be distant from the support and the load application areas; hence, Saint Venant\u2019s criterion is satisfied [30]. In order to develop a virtual instrument, the NI Measurements and Automation Explorer (MAX) software was used. This is a module embedded into software from National Instrument (NI) that allows access to all devices and NI systems. Within this software, communication between NI 9219 module and carrier USB NI 9171 is set up. After the input module is registered, a virtual instrument (VI) is developed in LabVIEW\u00ae Academic Suite" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002408_978-3-319-99522-9_9-Figure9.24-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002408_978-3-319-99522-9_9-Figure9.24-1.png", + "caption": "Fig. 9.24 The 3-RPS parallel manipulator", + "texts": [ + " The extensible legs connect the moving platform by spherical joints and the base by means of revolute joints, their axes being parallel to the opposite edge. Each leg is made up of a cylinder and a piston connected together by a prismatic joint. Three actuated revolute joints of the legs drive this spatial manipulator. The frame Ox0y0z0 with its z0 vertical axis is attached at the centre O of a fixed base. The platform is initially located at a central configuration, since the centre G of the frame GxGyGzG is fixed at OG \u00bc h elevation (Fig. 9.24). Six variables, namely three coordinates x G 0 ; y G 0 ; z G 0 of the mass center and three classical Euler angles a1, a2, a3 can describe the absolute position and orientation of the moving platform in the fixed reference frame. However, the mechanism is only a 3-DOF device; therefore, only three of them are independent. Here, z G 0 , a1, a2 will be chosen as independent variables and x G 0 ; y G 0 , a3 are parameters of parasitic motions. The three parasitic motions from the six commonly-known motions of the platform are permanently dependent on three independent variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002114_aqtr.2018.8402733-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002114_aqtr.2018.8402733-Figure1-1.png", + "caption": "Fig. 1. VTOL geometric model", + "texts": [ + "\u2211N i=1 ( ~Fi \u2212 ~\u0307Pi )T \u00b7 \u2202~vi\u2202q\u0307n + \u2211N i=1 ( ~Ti \u2212 ~\u0307Li )T \u00b7 \u2202~\u03c9i \u2202q\u0307n = 0 (5) This is a system of n equations fully describing the motion of the system of rigid bodies. All the forces, moments and velocities are expressed in world frame. Therefore, in the following we are interested in expressing the linear and angular velocity and momentum as well as the applied forces and torques in the world frame using the generalized coordinates q = [q1, . . . , qn]T C. The geometric model As can be seen in Figure 1, the proposed vehicle is formed out of three rigid bodies: 1) The main body containing two electric motors, the chassis of the vehicle and a semicircular capsule. 2) The front right arm with two motors 3) The front left arm with two motors The front arms are connected to the main body by one degree of freedom joints, thus each arm is allowed to rotate relative to the main body about the \u2212\u2192 j1 axis. Let {\u2212\u2192 i0 , \u2212\u2192 j0 , \u2212\u2192 k0 } denote the versors of the world coordi- nate system, {\u2212\u2192 i1 , \u2212\u2192 j1 , \u2212\u2192 k1 } the versors of a coordinate system attached to the main body of the vehicle and {\u2212\u2192 i2 , \u2212\u2192 j2 , \u2212\u2192 k2 } be a coordinate system attached to the right arm respectively {\u2212\u2192 i3 , \u2212\u2192 j3 , \u2212\u2192 k3 } be a coordinate system attached to the left arm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003825_1754337119831107-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003825_1754337119831107-Figure7-1.png", + "caption": "Figure 7. Arrangements for measurement of pressure distribution around 4.5 mm calibre pellets (numbers 1, 2, 3, etc., are the locations of pressure tapping holes; letters a, b, c, etc., are the area \u2018bands\u2019between pairs of tapping holes).", + "texts": [ + " Thus, the arrangement of Figure 6 allowed the tare drag on the dummy axial sting to be measured and subtracted from the overall (gross) drag, resulting in the required value of net drag force acting on just the pellet. This pellet overall net drag force was then used to calculate the pellet drag coefficient as shown in equation (5) Cd = drag force 1 2 rU2A \u00f05\u00de where drag force is the net drag force and A is the pellet cross-sectional area at the pellet head rim. Measurement of pressure distribution around the pellets The pressure distribution on the surface of the four pellets numbered 2 to 5 in Figure 2 was measured as illustrated in Figure 7. A hole of 0.2mm diameter (item A) was drilled on the pellet surface so that the hole formed a pressure tapping that communicated with the internal cavity of the pellet. The pellet was then secured onto a flanged stainless steel tube (item B in Figure 7) and sealed using instant adhesive. The tube was then used as a sting to support the pellet with its nose at 10mm downstream of the wind tunnel nozzle exit. The other end of the tube was connected to a pressure transducer using flexible small-bore tubing. In this way, pressure was measured at 18 tapping locations around the pellet, numbered 1 to 18 in Figure 7. A new identical pellet was used for each new tapping hole location, with all 18 pellets purchased in a single batch. This set of 18 surface pressure measurements was supplemented by an additional set of 10 measurements of pellet base pressure, made at radial locations numbered 19 to 28 in Figure 7. These 10 measurements of base pressure distribution were made with a micro-pitot tube (0.45mm i.d. and 0.80mm o.d.) positioned ;0.2mm downstream of the pellet base plane on a radial traverse system. The 28 surface pressure measurements around the pellet were also used to analyse the drag distribution around the pellet by means of pressure\u2013area numerical integration. The surface of the pellet and its base were divided into a number of surface area \u2018bands\u2019 as shown in Figure 7. The area bands are labelled with letters \u2018a\u2019, \u2018b\u2019, and \u2018c\u2019 etc., in order to distinguish them from the pressure tapping locations labelled with numbers 1, 2, and 3 etc. Each surface area band lay between two pressure tapping holes or two base pressure measurements in the case of the pellet base. As an example, the inset schematic in Figure 7 shows that the area band labelled \u2018d\u2019 (on the front face of the pellet) lay between the two pressure tappings 4 and 5. The pressure, pd, which acted on the surface of area band \u2018d\u2019, was considered to be the average of the pressures measured at tappings 4 and 5, that is, pd = (p4 +p5)=2. The pressure force that acted in the normal direction to surface area band \u2018d\u2019 was Fd = (pd pa)Ad, where Ad was the area of pellet surface band \u2018d\u2019. The force, Fd, for band \u2018d\u2019 was then resolved in the pellet axial direction in order to calculate the contribution of band \u2018d\u2019 to the overall pellet drag. Denoting the resolution of force, Fd, in the axial direction as Fdx, this resolved force was calculated, most conveniently, using the resolved area of band \u2018d\u2019 in the axial direction, that is, p(r25 r24) (see Figure 7). Therefore, as shown in equation (6), Fdx is Fdx = p4 +p5 2 pa p r5 2 r4 2 \u00f06\u00de The calculated axial drag force Fdx from band \u2018d\u2019 was a result of the mean pressure of band \u2018d\u2019, pd = (p4 +p5)=2, deviating from the atmospheric pressure pa. The axial force contributions of all 27 area bands around the pellet were then summed algebraically, which provided the overall pellet net drag force, FD. This overall net force was then entered as the drag force in equation (5) in order to calculate the pellet overall drag coefficient", + " The following factors were found to be crucial in achieving high repeatability and accuracy of Cd measurements: precise alignment of the pellet axis of symmetry with the axis of the jet; minimising pellet tare drag force, by shielding the load cell and pellet sting from the air jet; checking daily the calibrations of the data acquisition and load cell measuring systems; and measurement of local atmospheric pressure hourly, using a calibrated pressure transducer. Accuracy of Cd calculated from surface pressure distributions. Error propagation analysis was also used to estimate the error in overall pellet Cd obtained from surface pressure distribution measurements. The measurement of the radial location of each of the 28 pressure tappings was subject to an estimated random error of 0.05mm. As can be deduced from Figure 7, an error in the radial location of a shared tapping almost always changed the estimated relative area sizes of adjacent bands, but it did not change significantly the calculated overall projected area of the pellet. Rather, such an error introduced a random distortion of estimated drag force distribution around the pellet and an associated error in Cd. This error is not straightforward to estimate, but it is believed that its contribution to the overall error in Cd was small and was thus neglected", + "093, which is almost 10% greater than unity. Concluding this section, the measured values of Cd for spheres, obtained with the load cell and with pressure distribution methods, were in good agreement with published values at comparable values of free-stream Re and Ma. Pressure and drag force distributions around the dome-head pellet Figure 10 shows the measured distributions of pressure coefficient and drag force around the dome-head pellet. Pressure coefficients are shown for 27 surface area bands labelled \u2018a\u2019 to \u2018aa\u2019 (see Figure 7) around the dome-head pellet at Re ;54,000 and Ma ;0.58. For information, the data of Figure 10 have also been replotted in Figure 11, but this time, following convention, the x-axis shows the axial distance (streamwise direction) from the pellet nose tip. Figure 10 shows that the flow was brought to rest at the stagnation point at the centre of the pellet nose, and as a result, the pressure coefficient for band \u2018a\u2019 rose to just over 1.0. Further away from the stagnation point, close to the pellet head rim (area bands \u2018g\u2019 and \u2018h\u2019), the pressure coefficient was close to its lowest value of approximately 20" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001356_caidcd.2009.5375000-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001356_caidcd.2009.5375000-Figure1-1.png", + "caption": "Figure 1. Heat Distribution on the Contact Area", + "texts": [ + "00 \u00a92009 IEEE the calculation of gears temperature field, such as Wang Tong and Long Hui have used finite element method to analysis the bulk temperature of gears, and got a lot of useful the results and conclusions [3, 4]. However, it is very difficult to deal with the moving heat load in the finite element model, so the research on the real cause of the scuffing damage by the transient temperature has been very limited. 2.1. Moving Heat Load of the Gear Transient Temperature According to the well-known Block transient temperature calculation model, the width of contact interface is . Due to the relative sliding friction, the heat is generated in the contact area, as shown in Figure 1. While, the transient temperature rise (flash temperature) may be regarded as a heat generated by friction to the speed of movement along the contact plane 12b [5]. At this point, the transient temperature rise (flash temperature) may be regarded as the heat source generated by friction, has a movement along the contact plane in the speed of sq v , solving the temperature distribution of the gear. The distribution of the heat source and contact stresses is the same in the border: 2 1 max )(1 b xqqs 1 Where is instantaneous heat source intensity, is the maximum heat intensity, is semi-width of Hertzian contact band" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002920_poc.0000000000000176-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002920_poc.0000000000000176-Figure4-1.png", + "caption": "FIGURE 4. Microneedle array. This figure can be viewed online in color at www.poctjournal.com.", + "texts": [ + " The initial commercial version of TAP is intended for professional use; however, because it is easy to use, TAP has the potential to enable decentralized blood collection and testing in locations such as the patient's home. On the top of the device, shown in Figure 2, there is a green button and a fill indicator window. On the bottom of the device, there is a layer of hydrogel adhesive that seals to the skin and holds the device in place during use. The bottom view shown in Figure 3 displays the collection opening that interfaces with the skin, microneedle array, sample access port, and hydrogel adhesive. The microneedle array, shown in Figure 4, is arranged in a 7.5-mm diameter circular pattern with 30 needles around the circumference. In the current device, each of the microneedles is 1 mm long and 0.350 mm wide. TAP inserts and retracts the microneedle array by an internal insertion mechanism, which takes less than 1 second (Fig. 5). The force of microneedle insertion is provided by a snap dome, which achieves sufficient acceleration to overcome skin elasticity and successfully pierce the skin. The internal device mechanism is Point of Care \u2022 Volume 17, Number 4, December 2018 ealth, Inc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000946_1.2936383-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000946_1.2936383-Figure2-1.png", + "caption": "Fig. 2 \u201ea\u2026 Mechanical gas face seal kinematic model in a flexibly mounted stator configuration; \u201eb\u2026 mechanical seal with the seal faces aligned showing the coning angle", + "texts": [ + " Estimation based on the Gauss\u2013 arkov stochastic models significantly reduces computational verhead by avoiding direct computation of the nonlinear Reyolds equation, while still being applicable to a wide range of perating conditions. Mechanical Gas Face Seal System. A mechanical gas face eal schematic is shown in Fig. 1. One rotating seal ring rotor is igidly mounted to the rotating shaft. A nonrotating seal ring staor is flexibly mounted to the housing by a spring support and an lastomeric o-ring, which also acts as the secondary seal. The seal inematic model is shown in Fig. 2 a . The X and Y axes are fixed n a plane parallel to the rotor face plane, and the Z axis is orinted along the shaft axis. Rotor misalignment is not considered n this paper; therefore, the rotor face is perpendicular to the shaft xis. The stator support flexibility allows three degrees of freeom: stator axial translation Z from its equilibrium clearance C0 and tilts about the X and Y axes X and Y, respectively . igure 2 b shows the schematic side view of the rotor and a oned stator with a coning angle " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000691_car.2009.92-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000691_car.2009.92-Figure4-1.png", + "caption": "Figure 4. Basic HEDPA hardware platform and key equipments", + "texts": [ + " Thereinto, the mother robot refers to the \u201cbig robot\u201d coincide with special conditions, such as having the ability of bearing, deploying and collecting the son robots (and sensor nodes), while the latter refers to the \u201csmall robot\u201d having the ability of bearing and deploying the sensor nodes. In Fig.2, considering the given requirements, the restriction of the environment, and the characteristics of different robots and sensor nodes, the mother robot may bear the son robots and the sensor nodes to enter the region where the sensor nodes should be deployed and the information should be collected. According the result of intelligent planning fulfilled by the mother robot, the remote sever, or the operator (see Fig.4), the cooperation of network nodes deployment and perception can be achieved in direct cooperation manner (similar to the traditional architecture), or/and with the hierarchical expanding cooperation, which means the mother robot deploying the son robot, and the son robot deploying the sensor nodes further. Comparing with the traditional architecture, the expanding of HEDPA has three-level meanings, which are the situation is expanded where the sensor nodes can be deployed , the deployment and perception procedure can be gradually carried out, and the deploying and perceiving performance of the whole cooperation system can be advanced iteratively", + " Considering the aforementioned constraint conditions, the tasks under HEDPA can be classified as follows. BDPT (Basic Deployment and Perception Task): The robot cooperated deployment and perception tasks satisfy all of the aforementioned constraint conditions. CDPT (Complex Deploying and Perception Task): The robot cooperated deployment and perception tasks other than BDPT. The following contents of this paper are mainly focused on BDPT of HEDPA. IV. SOME KEY ISSUES OF HEDPA The fundamental hardware platform of HEDPA is shown in Fig.4. Thereinto, the mother robot is rebuilded from PeopleBot robot, the son robots are rebuilded from e-puck robots. The mother robot is equipped with active vision, selfmade integrated controller of UHF (Ultra High Frequency) RFID (Radio Frequency Identification) and WSNs, wireless Ethernet and transceiver. The mother robot fulfills the hierarchical expanding deployment of the son robots and the self-made sensor nodes, then achieve the cooperation perception of the environment and the objects. Optimizing Procedure 1) Robot path planning combination optimal model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000012_6.2008-4509-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000012_6.2008-4509-Figure7-1.png", + "caption": "Figure 7- Rotating Brush Seal Assembly", + "texts": [], + "surrounding_texts": [ + "American Institute of Aeronautics and Astronautics\n092407\n6", + "American Institute of Aeronautics and Astronautics\n092407\n7", + "American Institute of Aeronautics and Astronautics\n092407\n8\nTable 1 shows the seal combinations tested at Rexnord Corp. Figures 31 thru 36 show the performance of several seal combinations. Testing was completed at Rexnord which resulted in the seals been tested to 12,000 rpm. The approximate seal inlet temperature was 80o F" + ] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure5-1.png", + "caption": "Fig. 5. Second metamorphic limb with phase e1", + "texts": [ + " Therefore, phase e1 of metamorphic joint conform motion type of submanifold T (3) \u00b7U(p2,w2,v2) and phase e2 of metamorphic joint coincide motion type of submanifold T (3) \u00b7U(q2,u2,v2) combining the above constraint conditions. Therefore, transformation of motion type from submanifold T (3) \u00b7U(p2,w2,v2) to submanifold T (3) \u00b7U(q2,u2,v2) is correspond with transformation from phase e1 to phase e2 of metamorphic joint. 2) Type synthesis of second metamorphic serial limb: Submanifold T (3) \u00b7U(p2,w2,v2) limb has been generated by R(p21,v2) and R(p22,w2) \u00b7R(p23,w2) \u00b7R(p24,w2) \u00b7T (u2) with respect to local coordinate system o2 \u2212 u2v2w2. The metamorphic limb is synthesized with the phase e1 in Fig. 5. Link a is fixed to base, and rotational joint of axis 1 is locked. Axis 2 rotate along with axis v2 direction. Axis 3 rotate along with axis w2 direction. Axis 4 and axis 5 are parallel to axis 3 with w2 direction, end output link translate along with axis 6. Then metamorphic limb takes on mobility five which can realize motion type of submanifold T (3) \u00b7U(p2,w2,v2). Keeping rotational joint of axis 1locked, rotate link c to axis 3 parallel with axis u2 direction in Fig. 6. Axis 2 rotate along with axis v2 direction, axis 4 and axis 5 are parallel to axis 3 with u2 direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000352_cae.20144-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000352_cae.20144-Figure4-1.png", + "caption": "Figure 4 Toy car major parts.", + "texts": [ + " A robot control unit is therefore re-programmed including pallet presence control interlocks. This has eliminated malfunctions due to the misalignment of pallets in buffers, consequently improving the availability of the robot assembly station. A simple toy car is chosen as a production model. The model together with constituting parts was first built on computer using SOLIDWORKS-2004 software in 3D as shown in Figure 3. This model has five parts; a main body, a hat, four wheels, four nuts, and assembly base which are shown in Figure 4. Designed model has been produced in CNC machine tools. The parts placed on existing pallets are stored in the AS-RS shelves. The CIM software employed in this project is COSIMIR CONTROL by FESTO GmbH. It is a highly packaged, stand-alone cell control development, integration and run-time software tool kit. Simulation was carried out on the COSIMIR Educational/Professional software. The link between the software simulation and integrated hardware system is also provided by COSIMIR Professional. The Robot assembly unit is trained for assembling by hand-hand unit and fine tuned off-line separately by COSIROB" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000294_14399776.2008.10785987-Figure23-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000294_14399776.2008.10785987-Figure23-1.png", + "caption": "Fig. 23: Electrospindle radial displacement on motor side measuring plane, at different rotational speeds", + "texts": [ + " Figures 21a and 21b show experimental radial and axial load capacity on the nose respectively for different bearing supply pressures. Fig. 21a: Radial stiffness on the nose Figure 22 shows a similar spindle driven by an asynchronous motor with closed loop speed control, see Belforte et al. (2007D). The nominal power is 2.5 kW and speed is 75000 rpm. A tool can be mounted to test the electrospindle during the machining process. Motor and discharge air temperatures are controlled by a closed cooling circuit. Rotor orbits on the motor side measuring plane at different rotational speeds are shown in Fig. 23. They are measured with the same capacitive displacement transducers used for the pneumatic spindle. 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 8: 20 3 0 Se pt em be r 20 15 52 International Journal of Fluid Power 9 (2008) No. 3 pp. 45-53 Other studies are performed on applications where high stiffness is not required, but whirl stability must be improved. In particular, a spindle of diameter 37 mm with air bearings and pneumatic turbine was designed for textile applications, as is visible in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003418_ieem.2018.8607695-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003418_ieem.2018.8607695-Figure5-1.png", + "caption": "Figure 5: + # $,", + "texts": [ + " At the end of the layer, z for the current layer is incremented of the layer thickness and the process loop at the layer deposition stage. To validate the abacus based simulation, two case studies have been carried out. The first one to compare the results to a finite element simulation and the second one to compare the results to a built part. For the first case study, the results obtained using Abacus simulation have been compared to results from finite element simulation as proposed by [9], [10] and shown fig.5. The simulation represents 4 mm of beam path or 7.5 ms of built. The FE simulation has lasted 5h on an Intel Core i7 3GHz PC mounted with 8Go of RAM. On the same computer, the abacus based simulation duration was 5 min 30 s. The temperature distribution along a line swept by the beam is shown fig. 6. In this case the beam has traveled from X=0 to X=4 mm and the picture is taken when the beam is at X=4mm. In this figure limited difference between FE considered as the reference and abacus that needs to prove its reliability can be observed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003470_amcon.2018.8614963-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003470_amcon.2018.8614963-Figure4-1.png", + "caption": "Fig. 4 Test workpiece with inclined plane and text machining features", + "texts": [ + " Then, tool data, machining parameters and path data from Mastercam NCI were used and mainly sorted in accordance with Gcode-1001 and 20004 to be output to the Excel template. In addition, screenshot of the actual Mastercam verification could be used, where the material picture was named as stock and the coordinate system as G54, while each process name started sequentially from 1 to the end. The proposed methodology was implemented in a window-based environment using the Visual C# programming language. A semi-sphere workpiece with inclined plane and text machining features (Fig. 4) is used as an example and the NCI tool path is generated by the Mastercam software. Figure 5 displays the execution dialogue of the developed interface where the NCI file is converted into the NC program. Moreover, the NCI file can also be transformed to the process operation sheet indicating tooling, raw material and machining parameters, as shown in Fig. 6. A solid cutting simulation software package, VERICUT\u00ae is used to confirm the generated NC data. Given the raw material size, the specifications of the cutting tool, NC data, the type of controller, and the kinematics chain of an NC machine tool, it can interactively simulate the material removal process of NC data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000551_icara.2000.4803997-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000551_icara.2000.4803997-Figure1-1.png", + "caption": "Fig. 1. Setup of robot with nine IR sensors with a narrow infrared beam and three omnidirectional wheels. With these the robot can rotate and drive in arbitrary direction.", + "texts": [ + " The accuracy of this information depends on the sensor and whether it is measured relatively or absolutely. If relative sensors like rotary encoders, inertia or velocity sensors are used, additional absolute position information must be known to put the local data into relation with the global position. Over time, if no absolute sensor information is available, the positioning error increases. In our setup there are three sensor systems available to determine the position. The relative position is measured with rotary encoders that are mounted on the three omnidirectional wheels (see Fig. 1). In addition the robot is equipped with nine infrared distance sensors. These sensors, also shown in Fig. 1 with their narrow infrared beams, have a range of roughly up to 50 centimeters. They have nonlinear sensor characteristics and their sensitivities change with respect to the measured distances. Especially small obstacles can be easily missed, because the IR beam only detects larger objects that intersect the small band (see Fig. 1). The third sensor system here is called indoor GPS. Two transmitters are located at fixed positions at the ceiling. By a receiver that is mounted on each robot the actual absolute position is calculated. The accuracy of this system strongly varies dependent on the actual position. The experimental setup is situated indoors in a spatially confined area that is defined as the mission space MMission. In the mission space three identical robots are acting and trying to estimate their position. A map of obstacles in the mission space is available to make use of the IR sensors for position estimation", + " To get the probability grid of the odometry MOD(k), the previous overall grid Mres(k \u2212 1) is convoluted according to (5) with the kernel KOD: MOD(k, x, y, \u03d5) =\u2211 d\u03d5 \u2211 dy \u2211 dx Mres(k \u2212 1, x+ dx y + dy \u03d5+ d\u03d5 ) \u00b7KOD(k, dx dy d\u03d5 ). (8) The driven distance \u2206p is considered as the eccentricity of the kernel and the shift of dx, dy and d\u03d5 of its origin. For the angle the shifting is periodical. In the x- and y-axis the parts shifted to the outside of the map are ignored, and the parts shifted to the inside are set to zero. The result is again a probability grid that is dependent on the odometry and the previous measurement that was available as probability grid (see Fig. 6). C. Infrared distance sensors As can be seen in Fig. 1 each robot is equipped with nine infrared sensors. These sensors measure the radial distance to an obstacle. If the obstacles are associated with positional information, these sensors can be used to generate additional position information. To associate the positional information, introduced as feature grid Mfeat, the map of the mission space Mmission is transformed into an occupancy grid Mocc. An occupancy grid is a discrete map that consists of cells containing a probability that this cell is occupied by an obstacle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002787_ipec.2018.8507407-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002787_ipec.2018.8507407-Figure3-1.png", + "caption": "Fig. 3. Fabricated stator and rotor shaft.", + "texts": [ + " In addition, we propose much more convenient method to identify the current-force factor and the torque constant from the back EMF. Equation (7) shows the qaxis voltage at open circuit; d- and q-axis currents are equal to 0 A. zzkv zeeeeq '\u03a8\u03c9\u03a8\u03c9\u03a8\u03c9\u03a8\u03c9 \u03c9\u03c9\u03c9\u03c9 +=+= (7) In (7), the coefficients \u03a8z \u2019 and \u03a8\u03c9 are included, as a result, the current-force factor and the torque constant can be identified from the measurement of back EMF without measurements of the current, torque, and active axial force. IV. EXPERIMENTS Figure 3 shows the fabricated stator and the rotor shaft. In Fig. 3(a), stator cores are composed of laminated silicon steels. The three-phase winding is installed in the center stator core. The number of turns is 90 per tooth. In Fig. 3(b), the rare-earth material is used in the three-layer permanent magnets and passive magnetic bearings. The rotor outer diameter is 27 mm. The blue part is just plastic material to hold the segment permanent magnets. Figure 4 shows the active axial force with respect to daxis current. This is conventional method to identify the current-force factor. The active axial force is proportional to d-axis current in desaturation region, and then, the rate of change of the force is just the current-force factor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.38-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.38-1.png", + "caption": "FIGURE 16.38 Force equilibrium in the transverse direction of a tailored weld.", + "texts": [ + " The limiting thickness ratio (LTR) (for same materials with different thicknesses) or limiting strength ratio (LSR) (for different materials with a prespecified thickness ratio) is the thickness or strength ratio at which one material just reaches initial yield strength when the other material reaches its forming limit. When this value is exceeded, failure may occur parallel to the weld, with cracking occurring on the side of the material of lower thickness or strength. When the thickness and strength ratios are less than the LTR and LSR, respectively, plastic deformation would occur in both materials. To develop an expression for the LTR or LSR, let us consider Fig. 16.38, which shows two materials (1 and 2) that are joined by a laser weld. Neglecting friction effects, equilibrium in the transverse direction gives F1 = F2 (16.40) where F is the force per unit width and the subscripts 1 and 2 represent the materials 1 and 2, respectively. Equation (16.40) can be written as \u03c31h1 = \u03c32h2 (16.41) where \u03c3 is the true stress and h is the instantaneous material thickness. The true strain, \u03b5z, in the thickness direction is given by \u03b5z = ln ( h h0 ) (16.42) where h0 is the original thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003512_rcar.2018.8621755-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003512_rcar.2018.8621755-Figure3-1.png", + "caption": "Fig. 3 Sketch and components of the slave catheter/guidewire driving mechanism", + "texts": [ + " The surgeons make decisions based on the guidewire\u2019s situation in the vessels. So, the fluoroscopy image guidance is necessary in the CVI robotic system design. Moreover, as the medical robotic system, the concerns of the systematic design are listed as follows [13]: (1) Stable and safe motion of the guidewire in the vascular (2) Different speeds of the guidewire movement and rotation in the vascular (3) Easiness of CVI sterilization (4) Safety of the robotic-assisted surgery. The design of the SMN (Fig. 3 Sketch and components of the slave catheter/guidewire driving mechanism) adopted the modular design. The mechanism is composed of four modules: 1. Guidewire translation module 2. Guidewire rotation module 3. Force feedback module 4. Support module. Guidewire translation module and guidewire rotation module achieve the functions of the guidewire translation and rotation. The guidewire rotation module is fixed on a shifting board and the guidewire is clamped by a chuck in the front of the guidewire rotation module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001646_s12249-018-0983-6-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001646_s12249-018-0983-6-Figure5-1.png", + "caption": "Fig. 5. Compatibility assessment of mAb and SC towards five commercially available PFS models using AgNPs as sensor. a Plasmonic absorption results of mAb adsorption over GS long, GS short, and BD Bio PFSs and SC adsorption over GS long, GS short, BD Bio, BD Baked, and BD SCF PFSs. Adsorption was considered positive if SPR peak maximum \u2260 400 nm, whereas asdorption was considered negative if SPR peak maximum= 400 nm; b placebo controls for each protein-PFS combination; and c silicone oil-free PFSs controls. PFS, coated PFS (normal condition); uc-PFS, uncoated PFS; mAb uc-PFS, treated and mAb-exposed PFS, SC uc-PFS, treated and SC-exposed PFS", + "texts": [ + " On the other hand, shifted colors including orange, red, purple, brown, and clear (precipitated) solutions denoted adsorbed protein up to 1 \u03bcg onto the particular PFS. Finally, bright deep yellow solutions indicated adsorbed protein higher than 1 \u03bcg. As can be concluded, the color palette exhibited among the tested replicates revealed intrinsic heterogeneity within PFS models towards mAb or SC adsorption. Regarding the SPR of AgNPs, protein adsorption occurring onto PFSs was determined qualitatively (Fig. 5). Results showed characteristic and distinctive profiles of the establishment of protein-AgNP complexes in PFSs positive for adsorption, defined by having a SPR maximum value at other than 400 nm. While PFSs negative for adsorption maintained the characteristic profile of AgNPs solution with a SPR maximum at 400 nm. mAb-exposed PFSs showed that 71, 74, and 77% of the assays were positive for protein adsorption on the GS long, BD Bio, and GS short syringes, respectively (Fig. 5a). This was revealed by a red shifted resonance band (410 or 420 nm) and by the appearance of additional bands up to 600 nm. On the other hand, SCexposed PFSs showed that 87, 88, 91, 96, and 99% of the assays were positive for protein adsorption on the BD Bio, BD SCF, GS long, GS short, and BD baked syringes, respectively (Fig. 5a). According to the obtained results, GS long, BD Bio, and BD SCF syringe models emerged as the most suitable options for mAb and SC drug products. However, syringes with highest mAb and SC adsorption (GS short and BD baked syringes) presented an increase of only 6 and 12% in the number of assays positive for protein adsorption, respectively, when compared with most suitable syringes. Considering the method sensitivity (i.e., nanogram level), it was surprising to observe that adsorption on 15% of the total number of tested PFSs (n = 714) for both protein models was negligible (below the nanogram level). It is worth to notice that these containers correspond to highperformance PFSs, in terms of preventing protein adsorption, whose characteristics can be of special industrial interest. This assessment was also supported by the absence of interaction between bare AgNP solution and each PFS model, with or without exposure to formulation excipients of each drug product, as confirmed by the maintenance of SPR maximum at 400 nm (Fig. 5b). In order to confirm that proteins adhered to the PFS are effectively transferred to AgNPs, PFSs were treated to remove their silicone coating and make available its bare glass surface. Under this condition, protein adsorption onto the unlubricated glass surface of PFSs due to electrostatic interactions is maximized (40). SPR analysis showed that treated PFSs exposed to the mAb and the SC were positive for protein adsorption, while AgNPs solution from treated and non-exposed, and nontreated non-exposed PFSs exhibited a SPR maximum at 400 nm (Fig. 5c). These results confirmed that even under adsorption-promoting conditions to glass surfaces, proteins transfer from PFSs to AgNPs. Ultimately, the observed adsorption differences highlight the need of a case-by-case assessment as they result from the interaction of a specific drug product (API in combination with formulation excipients) with a PFS (41) given its intra- and intermodel characteristics (e.g., silicone oil layer structure and thickness, the occurrence of silicone oil particles, glass delamination, oxygen permeability, residual tungsten, Fe ions from needle, and distinct interfacial tension, among others) (17)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002434_s1068798x18080087-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002434_s1068798x18080087-Figure2-1.png", + "caption": "Fig. 2. Input shaft of a conical gear [2].", + "texts": [ + "eywords: radial clearance, contact angle, contact deformation, radial and axial dynamic loads DOI: 10.3103/S1068798X18080087 A radial ball bearing of 0000 type may be used to withstand purely axial loads. As an illustration, we show the input shaft of a cylindrical gear in Fig. 1 (from [1]); the input shaft of a bevel gear in Fig. 2 (from [2]); and the supporting roller of a four-high stand in a high-speed rolling mill in Fig. 3 (from [3]). Each shaft is mounted on two radial roller bearings intended to withstand the radial load and one radial ball bearing absorbing the purely axial load. This is ensured by a hole drilled in the housing under the bearing, whose diameter exceeds that of the bearing\u2019s external race. The radial ball bearing is used in horizontal and vertical electrical machines, even with relatively large axial loads on the shaft created by the drive" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.2-1.png", + "caption": "Fig. 18.2 Bodies of more general shape", + "texts": [ + " Since the contact point has no vertical component of velocity in Fig. 18.1, the instantaneous centres of rotation O1, O2 of the two bodies must lie on the common normal, and in fact they must be located at the centre of curvature of the surface immediately adjacent to the contact point, since if this condition were not satisfied, material points just upstream of the contact point would move to locations above or below the stationary contact point. Notice that the bodies do not have to be cylinders\u2014only the local region at the contact must be cylindrical [see Fig. 18.2] and the instantaneous centres O1, O2 will generally move along the common normal as sliding and/or rolling proceeds. In these figures, the instantaneous motion can be described by the angular velocities \u03a91,\u03a92, and with the anticlockwise positive convention shown V1 = \u2212R1\u03a91 ; V2 = R2\u03a92. (18.2) Whenwe consider themore general kinematics of three-dimensional contact, we find that a further condition is needed to determine the reference frame for rigid-body motions. We again choose a frame such that the contact point is stationary and the common normal at the contact point is vertical, but the system still has one degree of freedom, represented by rigid-body relative rotation about the common normal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure3-1.png", + "caption": "FIGURE 3 Gear-shaft interference assembly resultant deformation- radial (left), circumferential (center), and", + "texts": [ + " The static structural analysis was carried out for the gear-shaft interference set at nominal, minimum, and maximum interference. X axis was oriented radially, Y axis denoted angular orientation, and Z axis was aligned along the centerline of the gear. FEA results are listed in Table 2. Figures 3 and 4 show the radial deformation along the X axis in the gear body for nominal interference. The deformation due to interference is maximum at near the interface and reduces slightly with increase in diameter. The maximum radial deformation increases with increase in interference. Figure 3 (center) shows the circumferential deformation in the gear body for nominal interference. The circumferential deformation due to interference is small and varies slightly from one end to the other. The maximum deformation increases with increase in interference. \u00a9 2018 SAE International. All Rights Reserved. Figure 3 (right) shows the axial deformation in the gear body for nominal interference. The axial deformation due to interference is small. The maximum deformation shows slight increase with increase in interference. Overall, the gear body shows a significant radial increase in diameter with minimal changes in circumferential and axial directions. The effects of the tooth distortion on the various aspects of the assembled gear are discussed further. ISO 1328 [1] and ISO 10064 [2] give gear tooth profile and gear tooth helix measurement procedures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002514_msf.931.3-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002514_msf.931.3-Figure1-1.png", + "caption": "Fig. 1. Axisymmetric shell under the action of an asymmetric load. Displacements and resulting internal forces", + "texts": [ + "ntroduction The essence of the semi-analytic method is that the three-dimensional problem can be reduced to a two-dimensional by decomposition of the components of the surface load into Fourier series. So, if the load components U, V, W, parallel to the displacements u, v, w, change as: ( ) ( ) ( )cos , cos , sin ,n n nU U s n W W s n V V s n\u03b8 \u03b8 \u03b8= = = (1) then the displacements will also be periodic functions [1]: ( ) ( ) ( )cos , cos , sin .n n nu u s n w w s n v v s n\u03b8 \u03b8 \u03b8= = = (2) We will use an axially symmetric finite element in the form of a truncated cone. The directions of displacements and the resulting internal forces in the element are 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 Ltd, www.scientific.net. (#508824771, Rutgers University Libraries, Piscataway, USA-07/08/19,07:10:15) The Derivation of the Resolving Equations For the functions un(s), wn(s), vn(s), we take the following approximation: ( ) ( ) ( )2 3 0 1 2 3 4 5 6 7; ; ,n n nu s s w s s s s v s s\u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1= + = + + + = + (3) where i\u03b1 \u2013 some coefficients, determined from the boundary conditions at the ends of the element" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002156_012013-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002156_012013-Figure1-1.png", + "caption": "Figure 1. Geometry of single-axial groove bearing", + "texts": [ + " On the basis of previous studies, the influence of eccentricity and groove location on the oil film 2 1234567890\u2018\u2019\u201c\u201d pressure and bearing capacity of single-axial groove sleeve bearing is studied The two-dimensional Reynolds equation of incompressible fluid lubrication and finite long bearing can be expressed as: (1) where is the oil film thickness, is the oil film pressure, is the rotation speed of the sleeve, is lubricating oil dynamic viscosity, is the circumferential coordinate, is the axial coordinate. Geometry of single-axial groove bearing is shown in Figure 1, bearing has one groove along the axial direction, and the axial groove angle of the sleeve bearing is 36\u00b0. As shown in Figure 1, it is assumed that the groove is located in the position of A, B, C and D of the bearing. The oil film thickness without oil groove: , and the oil film thickness with oil groove: . is the radius clearance, is the eccentricity, is the attitude angle, is the depth of the oil groove, is circumferential coordinate . Under steady state condition, equation (1) can be reduced to (2) where is the non-dimensional oil film thickness: , is the non-dimensional oil film pressure, , is the width of bearing, is the length of bearing, is axial coordinate, ", + " (a) Position A (b) Position B 4 1234567890\u2018\u2019\u201c\u201d It can be seen from Figure 2(a), (b), (c), (d), Figure 3(a), (b) and Figure 4 that the non-dimensional oil film pressure distribution and non-dimensional bearing capacity of position A and B are obviously different with common sleeve bearing, and the non-dimensional oil film pressure distribution and non-dimensional bearing capacity of position C and D are basically the same with the common sleeve bearing. According to the location of each groove in Figure 1, the non-dimensional oil film pressure distribution and non-dimensional bearing capacity remain unchanged when the groove C, D are located in the divergence area of oil film; when the groove A , B are located in the convergence area 5 1234567890\u2018\u2019\u201c\u201d of oil film, the non-dimensional oil film pressure distribution will change greatly. Compared to the sleeve bearing with groove located in the divergence area, the non-dimensional bearing capacity of the sleeve bearing with groove located in the convergence area will decrease to a certain extent, the amplitude of the decrease will change with the different position of groove in the convergence area of oil film" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.6-1.png", + "caption": "FIGURE 16.6 Schematic of a plasma suppression setup. (From Watson, M. N., Oakley, P. J., and Dawes, C. J., 1985, Metal Construction, pp. 25\u201328. Reproduced by permission, TWI Ltd.)", + "texts": [ + " By comparison, other gases tend to form a plasma that absorbs the incident beam energy, under these conditions. Another use of the shielding gas is protection of the focusing lens. For low power applications, say, up to 1500 W, the normal flow of shielding gas (about 10 L/min) for protecting the weld is also adequate to protect the lens, as long as the gas stream is coaxial with the laser beam. A typical nozzle diameter would be 3 mm. However, at higher powers, separate shielding or helium cross-flow that is provided by a plasma suppression device (Fig. 16.6) is often applied. The plasma suppression unit directs a high-velocity jet of inert gas, usually helium, at the top of the weld to suppress or blow away the plasma. Plasma suppression serves to protect the lens and also to enable the heat supplied to reach the weld. The final form of the weld bead is affected by the mode of application of the shielding gas. For example, when welding steel with a 10 kW laser while using an axial gas nozzle of diameter 2 mm, no keyhole is produced at low helium flow rates of say, less than 5 L/min, since most of the beam energy is then absorbed by the plasma" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001603_978-3-319-73204-6_46-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001603_978-3-319-73204-6_46-Figure4-1.png", + "caption": "Fig. 4. Drone project and project goals", + "texts": [ + " While there is potential to gather evidence through every step of the design lifecycle, it still requires students to connect \u2018failures\u2019 to theory and thus improve basic knowledge, as well as reflect on deeper problemsolving skills. When designing the main project for the CEE program, the selected project and the project outcome must provide adequate coverage of multi-disciplinary features and provide project integration across courses. The current main project is to conceive, design, construct and test a drone using a 3D printer (Fig. 4). A drone has sufficient coverage of mechanical, electrical and program logic control components embedded in the design. These features of drones provide an ideal platform for solving open-ended problems and provide the students with independence to decide the extent of work they intend to conduct to achieve the project goals. Project appraisal considers the following guidelines: \u2022 Percentage of drone assembly designed and built using the 3D printer. \u2022 Degree of innovation and originality in the design" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000361_detc2009-87339-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000361_detc2009-87339-Figure2-1.png", + "caption": "Figure 2. SKETCH OF THE LUBRICATED JOURNAL BEARING AND THE TURBOCHARGER SHAFT WITH EXPLANATIONS OF THE USED VARIABLES.", + "texts": [ + " In the case of the turbocharger the shaft is supported by oil lubricated journal bearings which exert the forces Fb on the shaft and which are derived in the next section. Furthermore, the rotor is subject to gravity load Fg which is also added to the equation of motion: Mx\u0308 +(\u03c9G+C)x\u0307+ Kx = Fb(x, x\u0307)+ Fg + \u03c92Funb cos(\u03c9t). (12) The bearing forces are the sole source of nonlinearity in the model. The turbocharger shaft of radius R j is supported in oillubricated journal bearings of radius R and axial width W with a radial clearance cr = R\u2212R j (cf. Fig. 2). The position of the journal center in polar coordinates is e = \u221a u2 + v2 and \u03b3 = arctan( v u ). For small clearance-radius-ratio cr R \u226a 1 the oil film thickness is h = cr(1\u2212\u03bacos\u03d5) where \u03ba = e cr is the eccentricity ratio. To calculate the reaction forces of the bearing we consider a thin film flow between journal and bearing casing for which the pressure distribution inside the bearing can be described by Reynolds\u2019 equation [4]: \u2202\u03d5(h3\u2202\u03d5 p)+ R2\u2202z(h 3\u2202z p) = \u221212\u03c1\u03bdR2cr (( \u03b3\u0307\u2212 \u03c9 2 ) \u03basin\u03d5+ \u03ba\u0307cos\u03d5 ) , (13) where \u03bd is the kinematic viscosity of the oil, \u03c1 the oil density, p the pressure and \u03d5 the circumferential angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000968_tmag.2009.2012536-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000968_tmag.2009.2012536-Figure6-1.png", + "caption": "Fig. 6. Comparison of the distribution of magnetic flux density in the Z-axis direction of the overhang rotor.", + "texts": [ + " 1) IPM Motor With Overhang Rotor Model: This IPM motor model, shown in Fig. 5, consists of the four interior magnets, 24 slots stator, and the overhang rotor. The stack length of the rotor is 5 mm longer than the stator. The rotor and stator are constructed by laminating steel sheets. The thickness of the steel sheet is 0.5 mm and the space factor is 0.96. The influence of Table III shows the performance of the standard ICCG method and the AMGCG method. The AMGCG method shows better performance than the standard ICCG method. Fig. 6 shows the distribution of magnetic flux density in the Z-axis direction of the rotor. When compared with the result of the non-laminated rotor, we can confirm that the magnetic flux of 2) SPM Motor With Skewed Stator and Coils Model: This SPM motor model, shown in Fig. 7, has four poles and 12 slots. The thickness of the steel sheet in the stator is 0.5 mm and the space factor is 0.96. The stator and coils are skewed over 30 degrees. The skewed coils cause the magnetic flux in the lamination direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002358_gt2018-75432-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002358_gt2018-75432-Figure2-1.png", + "caption": "Figure 2: Mechanical Testing Specimen Plan used for the Study", + "texts": [ + " Volume fraction was performed using image analysis software provided by EDAX. Mechanical property testing of the buckets focused on two locations: airfoil and platform, primarily on stress rupture properties and LCF properties. Mechanical testing included 3 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 09/13/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use stress rupture testing at 1600\u00b0F/40 Ksi and isothermal LCF testing at 1650\u00b0F, 0.7% strain range. The specimens for mechanical testing are detailed in Figure 2. The stress rupture tests included two cylindrical stress rupture specimens from airfoil in the longitudinal direction, and one flat specimen from platform in the transverse direction, and the LCF tests included one cylindrical specimen from the airfoil in the longitudinal direction and another cylindrical specimen from the platform in the transverse direction. The locations of the specimens were obtained from the finite element models generated as a result of the studies published elsewhere [1]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000018_j.imagingsci.technol.(2008)52:6(060505)-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000018_j.imagingsci.technol.(2008)52:6(060505)-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of interactive touchdown developing method.", + "texts": [ + "n electrophotographic system is one of the most commonly used printing technologies during the past few decades. Demand on printing technology (e.g., printing speed, high quality, low cost, energy saving, and ecological technology) are increasing with rapid development of information technologies, and new machines having innovative and unique methods are being developed in many industries every year. An interactive touchdown developing method,1 which is shown in Figure 1, has characteristics of high print quality and high stability; i.e., it combines the advantages of both the two-component developing method and the monocomponent developing method. A two-component developer is carried by the magnet roller, and toner particles alone adhere on the surface of the development roll, forming a thin layer (two-component developing process). Then the toner particles are developed onto the latent images on the photoreceptor (monocomponent developing process). Thus, this method ensures excellent and long-lasting print quality" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000428_ssp.144.175-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000428_ssp.144.175-Figure3-1.png", + "caption": "Fig. 3. Solid Works Design of CNC Milling Machine", + "texts": [ + " Calculation of Drive Requirements on Feed Motor Shaft For the selection of an optimum servomotor and drive, it is necessary to find the drive requirements on the motor shaft. These are as below: a) Load Torque, b) Load Inertia, c) Maximum Speed. Generally, the following procedure is followed to find the load torque, load inertia and maximum speed requirements on the motor shaft and then it is selected proper motors in according to total inertia and maximum speed of table at the cutting process [3, 4, 5]. Main parts of CNC milling machine has been shown figure 2 and Fig. 3. It was used solid works to design the CNC machine as a detailed and It has been observed all movement the axes on the simulation process of the program before the producing the CNC machine. Some values can be taken from product catalogues, some of them can be taken from machine design or machine cutting books and some can be calculated by using concerning equations and constant values of machine tools and machining materials Torque for cutting at maximum speed of table T C = ( Ft.p B )/(2\u03c0 R\u03b7 ) =( 1100\u22c50" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000230_20080706-5-kr-1001.01206-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000230_20080706-5-kr-1001.01206-Figure1-1.png", + "caption": "Fig. 1. Definition of the camber angle", + "texts": [ + "3093 This paper is organized as follows: in the second section, the camber angle is quickly defined. In a third section, the developed measurement principle is detailed. In a fourth part, the feasibility of the measurement method is validated and its measurement precision is evaluated from real tests with the laboratory test vehicle. There are different definitions of camber angle. Here it is defined as the angle between the external plan of the wheel and a plan perpendicular to the road - considered as a plan - as shown in Fig. 1. The camber is defined as positive when the top of the tire is outside the vehicle body and negative in the opposite case. This definition has been selected because the other existing definitions take the tire torsion phenomenon under lateral solicitations into account. Indeed, this phenomenon can not be currently determined in real-time. Three major principles can be used to determine camber: the direct measurement, the indirect measurement or computation by measurement of other data and the use of maps issued from static tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001489_b978-0-12-812982-1.00002-3-Figure2.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001489_b978-0-12-812982-1.00002-3-Figure2.5-1.png", + "caption": "Fig. 2.5 Schematic view of the nozzle and the tapered hole resembling the key-hole. Arrows show the flow directions.", + "texts": [ + " Consequently, investigation into jet impingement onto the moving surface resembling the laser induced key-hole formation becomes necessary. The closed-form solution for the flow and temperature fields is very difficult due to the complex nature of the equations. However, the numerical solution for the problem is possible incorporating the appropriate boundary conditions. The mathematical arguments related to the flow and temperature field are given in relation to the previous study [12]. The geometric arrangement of the nozzle and the key-hole hole is shown in Fig. 2.5, while the nozzle and the hole configurations are given in Table 2.1. It should The steady and axisymmetric flow conditions are considered and the compressibility and variable properties of the working fluid are accommodated in the analysis. A constant temperature is considered at the hole wall to resemble the laser produced hole, that is, the hole wall temperature is kept at 1500 K as similar to the melting temperature of the substrate material. In addition, the hole wall is assumed to move with a velocity similar to the molten metal velocity as determined from the previous study [13]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003825_1754337119831107-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003825_1754337119831107-Figure1-1.png", + "caption": "Figure 1. Geometry and flow around a pellet. (a) Pellet cross section and geometry (CM is the pellet centre of mass and CP is the centre of pressure). (b) Main flow regions.", + "texts": [ + " The great majority of commercially available pellets are manufactured from soft pure or alloyed lead, while a few are made from harder metals such as tin, or partially from polymers or steel. Pellets come in a wide Department of Mechanical Engineering, University College London (UCL), London, UK Corresponding author: Nicos Ladommatos, Department of Mechanical Engineering, University College London (UCL), London WC1E 7JE, UK. Email: n.ladommatos@ucl.ac.uk variety of shapes, which range from that of a sphere to the common diabolo shape shown in Figure 1. The value of drag coefficient (Cd) for commercial 4.5mm calibre pellets varies widely, depending on pellet geometry and Mach number (Ma). For example, at Ma ;0.6, the Cd value varies between 0.3 and 0.8. By comparison, the Cd value for firearm bullets is generally smaller. At transonic speeds of Ma ;0.8 to 0.95, the Cd value for 5.56, 5.59 and 7.62mm calibre bullets ranges from about 0.2 to 0.3.15\u201317 Finally, the Cd value for 4.5mm smooth spheres at Ma ;0.6 is approximately 0.50.18,19 Air rifle and pistol pellets fly through the atmosphere usually at subsonic speeds, which range from Mach number of approximately 0.5\u20130.8. During pellet-free flight, flow separation often occurs at the head rim as shown in Figure 1, with the flow often reattaching close to the pellet tail rim near point C, but only to detach once again at the tail rim. The detaching boundary layers at the head and tail rims form free shear layers which engulf low-pressure recirculation zones labelled A and B in Figure 1(b). Separation of the flow from the pellet head and tail rims is an important contributory factor to pellet drag, as it causes lower pressures within the separation regions A and B, which contribute substantially to pellet drag. For example, the low-pressure recirculation region B, downstream of the pellet base, contributes a significant proportion of the total pellet drag (at least 35% at Ma ;0.6, see later). The pellet shown in Figure 1 has a fairly rounded front face so that the flow separating close to the head rim does so at an angle of approximately 45o to the pellet axis. In the case of pellets with a blunt flat face, the separating flow will leave the pellet rim at a greater angle of up to 90o, resulting in wider separation zones and wake and generally higher pellet drag. Pellets almost invariably have a hollow blunt base, and when they are in free flight, their large base area is exposed to the low-pressure recirculation region B (Figure 1). The design of the hollow base is deliberate. When the pellet is inside the rifle barrel, propelled by high-pressure gas, the gas pressure expands the base rim and forces it into the barrel rifling grooves, so as to seal the propelling gases. The barrel helical grooves force the pellet to spin, so that it leaves the barrel gyroscopically stabilised, similar to firearm bullets. The deep hollowing of the pellet shown in Figure 1(a) moves the pellet centre of mass (CM) forward so that it lies ahead of the centre of pressure (CP). This provides aerodynamic stabilisation, so that when the pellet leaves the barrel and assumes free flight, it is stabilised aerodynamically and gyroscopically. In contrast to air rifle pellets, which have blunt tails, firearm bullets generally have gradually tapered \u2018boat\u2019 tails, with the tail diameter gradually reducing towards their tail end to avoid separation of the flow from the side surfaces of the bullet tail" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001876_1464419318776742-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001876_1464419318776742-Figure2-1.png", + "caption": "Figure 2. (a) Bearing test rig, (b) dimensions of the tested bearing, and (c) normal and seeded bearing defect.", + "texts": [ + " The curve of AV values is obtained in between 0 and 1. A value of AV close to 1 indicates that the bearing is in normal condition and a value close to 0 indicates failure of the bearing AV \u00bc exp DI=cf \u00f06\u00de where cf is the scale factor. Experimental setup and data collection The proposed method for the performance degradation analysis of bearing is validated by acquired bearing datasets. Experimental setup is used to conduct a test on normal and seeded defects bearing. The bearing test rig details are given in Figure 2(a), which comprises a motor, speed controller, hydraulic loading system, and bearing housing and loading. For speed variations, this is associated with the synchronous motor, bearing set, coupling, and speed controller. The radial load applies through hydraulic jack with arm lever on bearing housing and the details are shown in Figure 2(a).The accelerometer is placed at a vertical position on the top and friction torque transducer fixed at the horizontal direction of the bearing housing. Photoelectric tachometer probe measures shaft speed. Friction torque transducer has two sets of beams with different stiffness to accommodate a wide range of experiment. The stiffer beams have side links affixed by machine screws to increase the transverse stiffness. The transverse stiffness (for one beam) is 2500 and 5700 lb/in., respectively", + " The bearing vibration signal is collected through piezoelectric accelerometer IMI 608 A11 model sensor and friction torque transducer mounted on bearing housing and eight channel National Instrument\u2019s PCI-4474 dynamic signal acquisition card. The DAQ control parameters are sampling frequency 5.12kHz, frequency lines 1600, and time duration 0.8 s. The bearing vibration data are captured and stored after every 10min and length of vibration signal for a period of 0.8 s is 4096 sample points. The standard 6205 ball bearings are tested and the geometric details and dimensions of the tested bearing are given in Figure 2(b). In this study, the run-to-failure test of bearings conducts with a different seeded defect and normal bearings are shown in Figure 2(c). The details of tested bearing information are given in Table 1. The bearing 2, test#2 (B2-2) and bearing 6, test#3 (B6-3) vibration signals in time domain under four different stages of degradation are depicted in Figure 3(a) and (b). The degradation stages are normal, slight, severe, and failure. The DIs for bearing 2, test#2 (B2-2) and bearing 6, test#3 (B6-3) are obtained after execution of steps discussed in \u2018\u2018Bearing degradation based on IMFs and K-medoids clustering\u2019\u2019 section form acquired signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002978_b978-0-12-409547-2.14488-9-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002978_b978-0-12-409547-2.14488-9-Figure2-1.png", + "caption": "Fig. 2 Simulated voltammograms showing the behavior observed for an organic molecule, for example, acetophenone, that undergoes a following chemical reaction upon oxidation. The electrode is a 5 mm radius platinum disk, and the rate constant for the homogeneous chemical reaction is 7.6 103 M 1 s 1.", + "texts": [ + " Measurements of this type are greatly facilitated by microelectrodes since solvents, such as alcohols or nitriles, that remain liquid over a wide temperature range can be used without catastrophic ohmic effects and the behavior of organic molecules such as o-nitrotoluene and nitrometisylene have been investigated. Adding or removing an electron from an organic molecule often creates a radical species that is not indefinitely stable and tends to react either with the parent molecule or other species in solution. For example, as illustrated in Eq. (1), electrochemical oxidation of ascorbic acid (vitamin C) produces an oxidized product that subsequently reacts with water to yield electrochemically inactive dehydroascorbic acid. \u00f01\u00de As shown in Fig. 2, the voltammetric response is significantly altered by the coupled chemical reaction and thus allows the energetics and dynamics of these homogeneous chemical reactions to be probed. This figure shows a case where the product of electron transfer undergoes a coupled chemical reaction. As the scan rate is increased, the contribution from the homogeneous reaction becomes less pronounced and the voltammogram approaches the shape of that for an electrochemically reversible process. The fact that ascorbic acid is electroactive means that its concentration in fruit juices can be determined without interferences from the coloration of the sample" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000182_09544062jmes1161-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000182_09544062jmes1161-Figure3-1.png", + "caption": "Fig. 3 A bearing pedestal flexibly supported by four triangular-hooked springs", + "texts": [ + " By substituting the corresponding mass, stiffness, and damping sub-matrices of equations (40) into the M, K, and C of equation (42), a 4k \u00d7 4k matrix containing the rotor\u2013bearing system characteristics equation (43) is obtained. This matrix has to be balanced prior to the eigenvalue calculations to avoid computational failure A = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 0 0 0 0 \u2212M\u22121(Kxx + KL \u2212 \u03b1 + Kxy) 0 0 \u2212M\u22121(Kyy + KL \u2212\u03b1 + Kyx) I 0 0 I \u2212M\u22121(Cxx + Cxy + \u03c9\u03b2) 0 0 \u2212M\u22121(Cxy \u2212 \u03c9\u03b2 + Cyy) \u23a4 \u23a5\u23a5\u23a6 4k\u00d74k (43) 3 EXPERIMENTAL PARAMETER ESTIMATION OF BEARING PEDESTAL 3.1 Bearing pedestal test rig The parameter identification is performed on the bearing pedestal illustrated in Fig. 3 without the stationary shaft. It consists of a 2.565 kg bearing pedestal that is flexibly supported by four triangular-hooked springs to realize a point and uncoupled effect to eliminate the cross-axis capabilities of the support system between the x- and y-directions. The spring preload in the ydirection can be adjusted by means of the stud and the pedestal housing is firmly secured to a rigid platform. The pedestal mass is excited about the horizontal position at a rotated angle of approximately 45\u25e6 by a stinger rod attached to a MB Dynamics Modal 50 electromagnetic shaker" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002588_icma.2018.8484720-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002588_icma.2018.8484720-Figure2-1.png", + "caption": "Fig. 2 The improved Circular Reference Self-calibration.", + "texts": [ + "00 \u00a92018 IEEE Proceedings of 2018 IEEE International Conference on Mechatronics and Automation August 5 - 8, Changchun, China The base of the robotic arm is cylindrical in the laboratory, the circular reference self-calibration method is not particularly suitable, while other calibration methods require a specific shape of the base. For the above situation, this paper proposes an improved circular benchmark self-calibration method. The improved calibration method only needs to randomly select three points that are not collinear on the side wall of the secondary manipulator (the end actuators of the main manipulator can moved to these three points) [9]. As mentioned above, this article selects three points on the robot arm as shown in Fig. 2. The four three points and their coordinates are ( )1 1 1 1, ,A x y z ( )2 2 2 2, ,A x y z ( )3 3 3 3, ,A x y z . Since the base is a cylinder, the three points 321 AAA of the side wall can be projected to the same plane, and the value of the z -axis direction does not have to be considered, and only the xoy plane can be considered to obtain the center position of the cylinder base. The coordinates of the middle point 3221 AAAA of the line segment 21 DD can be found as formula (1). 2,1, 2 , 2 ),( 11 =++== ++ n yyxx yxD nnnn dndnn (1) Lines 2211 EDED pass through points 21 DD , respectively, and are perpendicular to the straight line 3221 AAAA as formula (2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001813_978-981-10-8306-8_13-Figure13.10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001813_978-981-10-8306-8_13-Figure13.10-1.png", + "caption": "Fig. 13.10 The inverse Tresca p diagrams: a the yield surface and b the plastic strain increment surface", + "texts": [ + " However, their yield functions (and their averaging of summation or linear combination) provide diverse non-quadratic functions for incompressible, isotropic and symmetric elasto-plasticity. Also, their plastic strain increment functions (and their averaging of summation or linear combination) provide diverse non-quadratic functions for incompressible, isotropic and symmetric rigid-plasticity. The inverse Tresca p diagrams of the yield surface and the plastic strain increment surface are plotted in Fig. 13.10. The results in HW #12.2 and HW #13.9, suggest that Y = B = 3 K/2 and dY = dB = 2dK/3. Figure 13.10a confirms the former of these two relationships. As for deformation, the plastic strain increment for simple tension is unique as but this deformation is obtainable by diverse stress states between two pure shear stress states, (K, \u2212K, 0) and (K, 0, \u2212K). Meanwhile, the plastic strain increment for a balanced biaxial state, is unique as depBBTIII \u00f012 ; 12 ; 1\u00de \u00bc dB\u00f012 ; 12 ; 1\u00de but this deformation is also obtainable by diverse stress states between two pure shear stress states, (K, 0, \u2212K) and (0, K, \u2212K)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003676_978-3-030-12232-4_1-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003676_978-3-030-12232-4_1-Figure2-1.png", + "caption": "Fig. 2 Types of lubrication regimes present on the Stribeck curve", + "texts": [ + " In most cases, the friction and lubrication relationship is characterized with basis on \u03bcv F (oil viscosity \u00d7 sliding velocity/normal load) factor, in a diagram called Stribeck curve. This diagram summarizes the limits of hydrodynamic lubrication; see Fig. 1. Three zones can be identified, each one corresponding to a type of lubrication depending on the level of pressure established in the contact. For low pressure (0.1 to 50MPa), zone 1 corresponds to boundary lubrication; surface separation is ensured by lubricant molecules attached to the surfaces; see Fig. 2a. This type of lubrication is related to the physico-chemistry of surfaces and of lubricants, for low and moderate speeds and for relatively low loads. In zone 2, the hydrodynamic effect described by RE takes some importance and tends to separate the areas still in contact over a part of their asperities; this type of lubrication is the mixed lubrication; see Fig. 2b. Zone 3 corresponds to hydrodynamic lubrication and is described by RE; see Fig. 2c. In this region a full film separates the surfaces and friction is proportional to the speed if the lubricant viscosity is constant with temperature [48]. Notice that depending on lubrication regime, different surface interaction mechanisms occur, leading to distinct wear and friction responses. Friction behaviour in Stribeck curve is used to explain rubbing phenomena occurring in lubricated contacts. Transition between zones is explained as follows. For high values of \u03bcv F , friction coefficient is linearly ascending due to fluid film lubrication" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000012_6.2008-4509-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000012_6.2008-4509-Figure5-1.png", + "caption": "Figure 5-Split Ring Configuration", + "texts": [], + "surrounding_texts": [ + "American Institute of Aeronautics and Astronautics\n092407\n5\nThe test rig shown in Figures 2 and 3 has the following capability:\n\u2022 12,000 rpm \u2022 Delta Pressure Difference of 100 psid \u2022 Ambient Temperature Operation\nTesting was undertaken, using several seal configurations, rotational speeds, clearances and pressure\ncombinations on the following seals:\n\u2022 1thru 4 tooth labyrinth metal seals\n\u2022 1 split ring rotating brush seal\n\u2022 1 360 o ring rotating brush seal\n\u2022 2 split ring rotating brush seals arrange in tandem\n\u2022 2 360 o ring rotating brush seals arrange in tandem\n\u2022 2 split ring rotating brush seals arrange in tandem with centering springs\n\u2022 1 fractured carbon ring\nThe parts tested, with the exception of the carbon seals, are shown in Figures 4 thru 9.", + "American Institute of Aeronautics and Astronautics\n092407\n6", + "American Institute of Aeronautics and Astronautics\n092407\n7" + ] + }, + { + "image_filename": "designv11_92_0002580_022027-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002580_022027-Figure1-1.png", + "caption": "Figure 1. UCAV kinematic model parameter diagram", + "texts": [ + " With the help of neural network's nonlinear mapping ability and fault tolerance, the nonlinear mapping relationship between initial state, performance index (fitness function) and maneuvering flight control amount is constructed. And through the training completed network, a fast output of the flight control is achieved. In order to accurately describe the geometry of the maneuvering trajectory and the motion characteristics of the UCAV, a fine three degree of freedom motion dynamics model is established. The parameters of the model are defined as shown in figure 1. The flight control and the state variables are respectively defined as[ , , ]T\u03b1 \u03bc \u03b4 and[ , , , , , , ]Tx y h v m\u03b3 \u03c8 , the UCAV kinetic model can be expressed as[4]: cos sin cos cos sin cos sin cos sin cos cos sin sin cossin sin cos( sin )cos cos cos( sin )sin cos cos u x u y u h u x y h yx h u u u u u yx u u x v W y v W h v W T Dv g W W W m WW WL T g mv v v v v WWL T mv v \u03b3 \u03c8 \u03b3 \u03c8 \u03b3 \u03b1 \u03b3 \u03b3 \u03c8 \u03b3 \u03c8 \u03b3 \u03b3 \u03c8\u03b3 \u03c8 \u03b3\u03b1 \u03bc\u03b3 \u03b3 \u03c8\u03b1 \u03bc\u03c8 \u03b3 \u03b3 = + = + = + \u2212= \u2212 \u2212 \u2212 \u2212 += \u2212 + + \u2212 += + \u2212 2 2 max sin cos 1 1, ( , ), , 2 2 u u c u L u D v m cT T T v h L v SC D v SC \u03c8 \u03b3 \u03b4 \u03c1 \u03c1 = \u2212 = = = (1) In the formula: ( , , )x y h represents the position of UCAV in the inertial coordinate system; uv is the UCAV vacuum speed; ( , , )\u03b3 \u03c8 \u03bc represents the track inclination angle, track declination angle, track roll angle; \u03b1 is the angle of attack; m is the quality of UCAV; g is the acceleration of gravity; ( , , )T D L represents respectively Thrust, air resistance, lift; ( , , )x y hW W W and ( , , )x y zW W W represent the components along the axis for wind speed and wind acceleration; 93001" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003825_1754337119831107-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003825_1754337119831107-Figure5-1.png", + "caption": "Figure 5. Pellet drag force measuring system. (a) 17 mm \u00d8 nozzle. (b) Pellet. (c) Pellet sting (1 mm \u00d8 3 20 mm long + 1.5 mm \u00d8 curved support). (d) Shear-force load cell. (e) Air shield for load cell and sting (no contact with sting, except at bottom of load cell; side shields not shown for clarity). (f) Three-dimensional traverse (X, Y, and Z).", + "texts": [ + " The above values indicate that the turbulence intensity of the open tunnel was low. It compares favourably with values of turbulence intensity ranging from 0.30% to 2.5% at the centreline of open wind tunnels described in published sources.5,28\u201331,38\u201341 The low level of free-stream turbulence measured at Ma=0.58 and 10mm downstream of the nozzle exit plane suggests that turbulence did not have a significant influence on the measured values of drag and pressure coefficients presented in this article. The pellet drag was measured using the load\u2013cell system shown in Figure 5. The pellets were positioned with their nose 10mm downstream of the nozzle exit plane. The pellet support comprised two sections: the pellet sting itself (1mm o.d.3 20mm long) followed by a curved sting extension (1.5 mm o.d.) which was attached to a strain-gauged load cell that measured the pellet drag force. The shear load cell, normally used in precision balances of 1mg resolution, was insensitive to variations in the location of the point of loading and side loads. Part of the sting support and the load cell itself was shielded from the air flow so as to minimise the tare drag force on the pellet", + " Therefore, the force measured by the load cell, was the gross drag value, which included the air drag force on the pellet itself plus the \u2018tare\u2019 drag due to the air jet impinging on the sting. In order to evaluate the net drag on the pellet, the tare drag on the sting was measured (after each measurement of the gross drag), using the additional arrangement shown in Figure 6. An identical pellet was mounted, using a transverse sting (item D in Figure 6), to a support system (items E, F, G) which was independent of the load cell. An axial \u2018dummy\u2019 sting (item C), not connected to the pellet, was placed in position and attached to the load cell of Figure 5. Thus, the arrangement of Figure 6 allowed the tare drag on the dummy axial sting to be measured and subtracted from the overall (gross) drag, resulting in the required value of net drag force acting on just the pellet. This pellet overall net drag force was then used to calculate the pellet drag coefficient as shown in equation (5) Cd = drag force 1 2 rU2A \u00f05\u00de where drag force is the net drag force and A is the pellet cross-sectional area at the pellet head rim. Measurement of pressure distribution around the pellets The pressure distribution on the surface of the four pellets numbered 2 to 5 in Figure 2 was measured as illustrated in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001823_978-3-319-91262-2_52-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001823_978-3-319-91262-2_52-Figure4-1.png", + "caption": "Fig. 4. Two-Rotor Aero-dynamical System.", + "texts": [ + " The aim of this section is to implement the proposed methodology and verify its usefulness using real system. For this task, the Two-Rotor Aero-dynamical System (TRAS) is chosen. The TRAS is a laboratory set-up designed for control experiments. In certain aspects its behaviour resembles that of a helicopter. From the control and modelling point of view it exemplifies a high order nonlinear system with significant cross-couplings. The system is controlled from a PC. A schematic diagram of the laboratory set-up is shown in Fig. 4. The TRMS consists of a beam pivoted on its base in such a way that it can rotate freely both in thehorizontal and vertical planes. At both ends of the beam there are rotors (the main and tail rotors) driven by DC motors. A counterbalance arm with a weight at its end is fixed to the beam at the pivot. The state of the beam is described by four process variables: horizontal and vertical angles measured by position sensors fitted at the pivot, and two corresponding angular velocities. Two additional state variables are the angular velocities of the rotors, measured by tacho-generators coupled with the driving DC motors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001688_2018-01-0838-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001688_2018-01-0838-Figure1-1.png", + "caption": "FIGURE 1 (a) Journal bearing test rig (b) Location of pressure sensors", + "texts": [ + " The space between the bearing and shafts is filled with lubricating oil. In hydrodynamic lubrication technique, as the shaft starts to rotate, it tries to climbs the bearing surface and as rotational speed of the shaft is further increased, the lubricating oil forces itself into a wedge-shaped region. This wedging action of the lubricating oil helps in generating the pressure within the system. In the present work, in order to find out the pressure and frictional torque, a journal bearing rig TR-660 was used. Figure 1(a) represents the basic parts of the test rig and Figure 1(b) shows the location of pressure sensors. Journal bearing test rig is a sturdy versatile apparatus, easy to operate with provision to measure pressure at a different angular position on the bottom half of journal bearing. The journal is mounted horizontally on a shaft supported on self-aligned bearings, the shaft is rotated by a motor with timer belt. A loaded bronze flawless bearing freely slides over the journal and as it rotates bearing is formed, radial load is applied on bearing by pulling it upwards against journal by a loading lever", + " The rupture of the lubricating film at high speed (more than 1900 rpm) is the possible reason for this increase in the friction torque. This increase in friction torque is also made confirm by drawing the graph between rotational speed and maximum pressure generation in the oil film. Figure 4 represents the variation of maximum pressure generation in the lubricating oil film. The pressure values used in plotting the graph are detected by a pressure sensor number 7 located at 213.75\u00b0 from the lubricant oil entry point (see Figure 1(b)) on the bearing surface. It is seen that as the rotation speed is increasing, the oil film pressure also increases and reaches a maximum value of 1.15 MPa at 1600 rpm, and it falls to a minimum value (0.8 MPa) as the rotational speed is further increases to 1900 rpm. So, at 1900 rpm and 500 N load values, the oil film gets rupture and maximum pressure generated falls to 0.8 MPa. Also during the lower speeds, the sliding and rolling between the asperities of the journal and bearing will be very less, which would result in less friction torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001062_cimca.2008.135-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001062_cimca.2008.135-Figure1-1.png", + "caption": "Figure 1. The base-frame, location and a step movement of the platform in x-y plane.", + "texts": [], + "surrounding_texts": [ + "benchmark problem from the plane to the three dimensional real world [2], [3]. The kinematics and dynamics of the nonholonomic wheeled mobile robots have been introduced and developed by many researchers [4], [5], [6]. However, as stated in [1], none of them can be directly applied to a mobile inverted pendulum stabilization problem with the constrained steering angle of a nonholonomic platform. The system stabilized in [1] has no platform, and the free movement of the pendulum is only in forward-backward direction due to the system structure. This paper contributes the kinematics and dynamics of a nonholonomic mobile platform for the dynamic stabilization of a freely moving inverted pendulum in a horizontal planar path. The dynamic modeling of the system is accomplished using homogeneous coordinate transformations with the Lagrangian method. An intermediary joint and link are required to represent the two independent rotation modes of the rod [12], but the Denavit-Hartenberg convention [11] is not applied to keep the z-axis of the of the platform, intermediary and rod frames directed upward. This paper is organized as follows: Section 1 describes the four-wheeled mobile inverted pendulum system. Section 2 presents the main components and kinematics relations. The dynamic equations of motion are developed in Section 3. Section 4 presents the conclusion and future research directions on FWMIP systems and the mobile inverted pendulum. 2. Kinematics of the System We represented the two rear wheels by a single central fixed wheel, and the two steering wheels by a single virtual central steering wheel similar to [9], such that the system is simplified to a bicycle in x-y plane [4]. Virtual central steering angle \u03d5 provides a central steering radius r = lp cot(\u03d5). The right and the left steering wheel angles \u03d5r = atan(lp /(r+wp/2)), and \u03d5l = atan(lp /(r\u2013wp/2)) are set by a steering mechanism depending on the front wheel distance wp.\n978-0-7695-3514-2/08 $25.00 \u00a9 2008 IEEE DOI 10.1109/CIMCA.2008.135\n409", + "We assigned coordinate frames Fp and Fr, to the platform and to the inverted pendulum rod to accomplish the kinematics and dynamics analysis of the system as seen in Figures 1 and 2. The x-axis of the platform frame is oriented toward the front of the platform, and z-axis is selected upwards. Accordingly, the nonholonomic constraint restricts the rear wheel motion in local y direction. The frame is placed at the joint point of the rod to achieve easy stabilization of transversal rod motion, although it is known that placing the platform frame at the central fixed wheel simplifies the kinematic relation regarding this frame [4]. The position of the platform is described by the vector\nqp= (x, y, \u03c8), where x, y are the coordinates of the platform frame in the base frame, and \u03c8 is the orienta-\nMovement with a steering angle \u03d5 =0 gives a linear motion along the x-axis of the platform, while \u03d5 0 results in a tangential motion \u0394d of the platform on a circle with the steering radius r = lpC\u03d5 /S\u03d5. This circular path turns the platform \u0394\u03c8 = \u0394d / r amount about the z-axis with respect to the initial platform position. Rotations of a coordinate frames about x, y and z-axes and its translation to a point p = (x, y, z)T can be expressed by homogeneous transformation matrices\nrot(x,\u03b1) = 1 0 0 0 0 C S 0 0 S C 0 0 0 0 1 \u03b1 \u03b1 \u03b1 \u03b1 \u2212 ; rot(y,\u03b2)= C 0 S 0 0 1 0 0 S 0 C 0 0 0 0 1 \u03b2 \u03b2 \u03b2 \u03b2\u2212 ;\nrot(z,\u03c8)= C S 0 0 S C 0 0 0 0 1 0 0 0 0 1 \u03c8 \u03c8 \u03c8 \u03c8 \u2212 ; trans(x,y,z)= 1 0 0 0 1 0 0 0 1 0 0 0 1 x y z ; (1)\nwhere \u03b1, \u03b2, \u03c8 are angular displacements, and the abbreviations C\u03b8 and S\u03b8 denote cos(\u03b8 ) and sin(\u03b8 ) [11]. The nonholonomic motion of the platform is conveniently expressed by a chain of homogeneous translation and rotation transformation matrices. The final platform location shown in Figures 1 and 2 is obtained relative to the initial platform frame by the following chain of transformation matrices 0Tp1 trans(\u2013 lp , lp cot(\u03d5) , 0 ) . rot(z, \u0394\u03c8) . trans(lp , \u2013 lp cot(\u03d5) , 0) ; (2) Fp1 p1Tp2 Fp2 (3)\nFrom Fp2 to Fp1 the transformation p1Tp2 is p1Tp2= ( )\n( ) C S 0 S C C S S / S S C 0 C C S S C / S\n0 0 1 0 0 0 0 1\np\np\nl\nl \u03c8 \u03c8 \u03d5 \u03c8 \u03d5 \u03c8 \u03d5 \u03d5 \u03c8 \u03c8 \u03d5 \u03c8 \u03d5 \u03c8 \u03d5 \u03d5 \u0394 \u0394 \u0394 \u0394 \u0394 \u0394 \u0394 \u0394 \u2212 + \u2212 \u2212 \u2212 . (4)\nFurther, the initial platform frame is transferred to the base frame Fp0 by rotation and translation transformations. 0Tp1 = trans(lp , 0 , 0) . rot(z,\u03c8) (5) F0 = 0Tp1 Fp1 . The overall transformation 0Tp is obtained as 0Tp = 0Tp1 p1Tp2 .\n= C S 0 (C C +(S S )C /S ) S C 0 (S S (C C )C /S )\n0 0 1 0 0 0 0 1\np\np\nl x l y \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03d5 \u03d5 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03d5 \u03d5 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 \u2212 \u2212 \u2212 + \u2212 \u2212 \u2212 + .\n(6)", + "Thus, the movement (\u0394x, \u0394y, \u0394\u03c8) between the final and initial positions of the platform for the steering wheel angle \u03d5 and the tangential path distance \u0394d is \u0394\u03c8 = (\u0394d / lp ) tan(\u03d5 ) \u0394x = lp (C\u03c8+\u0394\u03c8 \u2013 C\u03c8 + (S\u03c8+\u0394\u03c8 \u2013 S\u03c8) / tan(\u03d5) ) , \u0394y = lp (S\u03c8+\u0394\u03c8 \u2013 S\u03c8 + (C\u03c8+\u0394\u03c8 \u2013 C\u03c8) / tan(\u03d5) ) . (7)\nThe steering angle \u03d5 =0 gives divide by zero in (7) due to approach of r to infinity. In that case (4) converges to trans(\u0394d, 0, 0). Consequently, the overall transformation comes out\n0Tp= C S 0 C S C 0 S 0 0 1 0 0 0 0 1 d x d x \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u2212 \u0394 + \u0394 + , (8)\nresulting in incremental kinematics relations \u0394\u03c8 = 0; \u0394x = \u0394d C\u03c8 ; \u0394y = \u0394d S\u03c8 . (9)\nThe rod frame Fr is placed to the free end of the inverted pendulum rod, so that the frame position conveniently indicates deviation from the vertical pose. At the platform-rod joint the rod rotates freely about y and x axes. For a rigorous dynamic representation of this two degree-of-freedom joint we use an intermediary link and joint that provides rotation about x-axis. pTri = rot(x,\u03b1) ; (10) riTr = rot(y, \u03b2) trans(0, 0, lr) (11) In this configuration, eliminating the virtual joint and link restricts the motion of the rod only into x-z plane. The transformation\npTr = pTri riTr (12)\nFp = pTr Fr (13) converts the coordinates from the rod frame Fr to the platform frame Fp. The transformation pTrv and riTr are\npTri = 1 0 0 0 0 C S 0 0 S C 0 0 0 0 0 \u03b1 \u03b1 \u03b1 \u03b1\u2212 ; riTr = C 0 S S 0 1 0 0 0 C C 0 0 0 1 r r l S l \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u2212 \u2212 ,\nand the combined transformation pTr is\npTr = C 0 S S S S C S C S C C S S C C C C\n0 0 0 0\nr\nr\nr\nl l l \u03b2 \u03b2 \u03b2 \u03b1 \u03b2 \u03b1 \u03b1 \u03b2 \u03b1 \u03b2 \u03b1 \u03b2 \u03b1 \u03b1 \u03b2 \u03b1 \u03b2 \u2212 \u2212 \u2212 \u2212 . (14)\nThe coordinate transformation matrices 0Tp , pTri and riTr represent the position and orientation of the platform and rod to derive the dynamics of the system.\n3. Dynamics of the System Let q =(q1, \u2026 qn)T be vector of generalized joint displacements of an n degree-of-freedom open chain mechanism. The derivatives of Lagrangian L= K P give the generalized joint-force \u03c4i at the joint-i.\n\u03c4i = dt d iq \u2202 \u2202 L \u2212 iq \u2202 \u2202 L , (15)\nwhere K and P are the overall kinetic energy and potential energy of the analyzed frictionless system [11]. Integrating kinetic energy of differential masses over the complete mass of the ith link gives the kinetic energy ki of the link. ki =\ni im\nd k\n= 12Trace[(\u03a3i p=1 \u03a3i r=1 0 i pq \u2202 \u2202 T Jpi( 0 i rq \u2202 \u2202 T )T qr qp )], (16) where Jpi = mi\nirdm (irdm)T dmi is the pseudo-inertia matrix of the ith link. The potential energy pi of the ith link is pi = \u2212 mi ga T 0Ti icmi , (17) where ga is the gravitational acceleration vector, mi is the mass, and 0Ti\nicmi = 0cmi is the center of mass of the ith link. The time and joint variable derivatives of the Lagrangian give the generalized jointforce\n\u03c4i = dt d ( ) iq \u2202 \u2212 \u2202 K P \u2212 ( ) iq \u2202 \u2212 \u2202 K P , (18)\n= \u03a3n r=i \u03a3n k=1 ( 0 r kq \u2202 \u2202 T Jpr( 0 r iq \u2202 \u2202 T )T) q k\n+ \u03a3n j=1\u03a3n k=1\u03a3n r=1Trace( 2 0 r k jq q \u2202 \u2202 \u2202 T Jpr( 0 r iq \u2202 \u2202 T )T)q k q j\n\u2212 \u03a3n r=i mr ga T 0 r iq \u2202 \u2202 T rcmr . (19)\nFurther, (19) can be put in Uicker-Kahn form [11].\n\u03c4i =\u03a3n j=1 Dij qj +\u03a3n j=1 \u03a3n k=1 Cijk q j q k + g i . (20)\nwhere\nDij =\u03a3 n r=max(i,j) Trace( 0 r jq \u2202 \u2202 T Jpr ( 0 r iq \u2202 \u2202 T )T) ;\nCijk =\u03a3n r=max(i,j,k)Trace(\n2 0 r j kq q \u2202 \u2202 \u2202 T Jpr ( 0 r iq \u2202 \u2202 T )T), and\ngi = \u2212 \u03a3n r=i mr ga T 0 r iq \u2202 \u2202 T rcmr , (21)\nare the inertial, coriolis, and gravitational terms of the equation of motion. The equation is further written in a matrix form" + ] + }, + { + "image_filename": "designv11_92_0003691_robio.2018.8665156-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003691_robio.2018.8665156-Figure3-1.png", + "caption": "Fig. 3. Mechanism of high-speed robot hand [15].", + "texts": [ + " The rest of this paper is organized as follows; Section II explains the human\u2013robot collaborative system we developed, Section III describes the proposed strategy for human\u2013 robot collaboration, Section IV discusses the stability and a collaborative error, Section V shows experimental results of collaborative motion and a collaborative peg-in-hole task and discusses the frame rate of the collaborative system, and section VI summarizes the conclusions obtained in this work. As shown in Fig. 2, our human\u2013robot collaborative system consists of \u2022 a high-speed robot hand (Section II-A), \u2022 an image-processing PC with a high-speed camera (Sec- tion II-B), \u2022 a real-time controller that receives the values of the state of the board (position and orientation) from the imageprocessing PC at 1 kHz and also controls the high-speed robot hand at 1 kHz, and \u2022 a board to be manipulated by the robot hand and a human subject (Section II-C). Fig. 3 shows the mechanism of the high-speed robot hand [15]. The high-speed robot hand has three fingers: a left thumb, an index finger, and a right thumb. Each finger has a top link and a root link, and the left and right thumbs rotate around a palm. Therefore, the index finger has two degrees of freedom (2-DOF), and both thumbs have 3-DOF. In addition, the robot hand has a wrist joint with 2-DOF (in Fig. 3, 1-DOF movement is illustrated). Thus, the hand has a total of 10-DOF in its movement. In order to measure the state of the board state (position and orientation), we used a high-speed camera called EoSens MC4086 produced by Mikrotron [16]. An image-processing PC equipped with a frame grabber board could acquire raw image data from the high-speed camera. The imageprocessing PC was also equipped with an Intel(R) Xeon(R) W5-1603 v3 2.8 GHz processor and 16 GB of RAM. The operating system of the image-processing PC was Windows 7 Professional (64-bit)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002228_kem.774.313-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002228_kem.774.313-Figure1-1.png", + "caption": "Fig. 1: Possible orientation of 3D printed specimens, a) scheme, b) actual examples", + "texts": [ + " The first stage of the experimental programme contained a comprehensive evaluation of microstructure, mechanical properties, fatigue durability and creep of the first batch of samples manufactured using the 3D printing technology, where the specimens were printed in the so called X position with specimen axis perpendicular to the built direction. The results were completed and discussed with analyses of damage mechanisms particularly using scanning electron microscopy. Experimental Material and Specimens As a result of the DMLS principle it is necessary to perform detailed evaluation of microstructure, mechanical, fatigue and other properties for all batches of specimens or components as a dependence of their orientation. There are five most important orientations schematically shown in Fig. 1, where z-axis represents the building direction layer by layer and x-y represent plane with the laser beam movement. In this work 0-X direction only was investigated. One of the most important factors affecting the printed material quality is input powder. Both types of powder, namely original one and already used sieved powder can be used for the technology. In this case, the original powder, manufactured by fluid spraying of Inconel 718 alloy was used for most of the specimens. Scanning electron microscopy (SEM) showed that the powder contained particles mostly of a globular shape of the diameter between 8 \u00b5m and 65 \u00b5m" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002923_icsgsc.2018.8541305-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002923_icsgsc.2018.8541305-Figure1-1.png", + "caption": "Figure 1. The illustration of the vehicle on the road", + "texts": [ + " In decoupling process, the three-phase mathematical model of PMSM will be changed into the two-phase mathematical model using Clarke and Parke transformations[3]. The Clarke transformation converts the balanced three-phase quantities into two-phase stationary reference frame. The Park transformation converts from stationary reference frame into a rotating reference frame. When PMSM is used as a driving force in an electric vehicle, this out of synchronization can cause problems. When the vehicle is moving on a flat road as shown in Fig. 1.(a), due to the effect of vehicle loads, there can be a magnetic slip, as shown in Fig. 1.(b). Although the direction of the flux is the same, with different speeds ( me ), the motor will be asynchronous. Not only on flat roads, the asynchronous condition also occurs in vehicles that are moving on uphill roads, as shown in Fig.1.(c). Because of the effects of gravity, where the vehicle gets as much force in the opposite direction, the vehicle can run backward. This condition is dangerous for the passenger of the vehicle. In Fig.1.(d), a magnetic slip condition is indicated when the vehicle is on an uphill road. Unlike when a vehicle moves on a flat road where the direction of the stator flux is equal to the rotor flux, the magnetic slip condition when the vehicle is running on an uphill has a stator flux direction that is not equal to the rotor flux. If the initial position of the rotor ( r ) differs from the controller's initial position, it can cause problems, especially if the PMSM is used as an electric vehicle, as illustrated in Fig. 1. This condition increases if the motor gets a large load. A large load on the motor can also cause problems. As explained by Harini, B.W. et al[4], the electric vehicle actually moves in the wrong direction when a large load that has reverse torque direction is given to the motor. Different from that paper where the system used a sensorless system, this paper will present the effect of initial position and load on PMSM if the system uses incremental encoder sensor. The initial position of the rotor is very important in the control of PMSM", + " et al have described an efficient method for identifying the initial position of a permanent magnet synchronous motor (PMSM) with an incremental encoder, even when a brake and/or a constant load torque is being applied[5]. However, in this paper, there is no explanation of the effect of the initial position on the motor current. Lei, Wang et al have proposed a method to identify the rotor position polarity[6]. This paper didn\u2019t describe the effect of the initial position on the motor current. All papers didn't research the phenomena as illustrated in Fig.1. The control system that will be analyzed uses Proportional Integral (PI) as a current controller and 95 978-1-5386-8102-2/18/$31.00 \u00a92018 IEEE incremental encoder as a sensor. The effect of the load and initial position on PMSM control will be analyzed on both road condition, as illustrated in Fig.1. II. PMSM MODEL The Park transformation converts from stationary reference frame into a rotating reference (d, q, 0) frame using (1). s s rr rr sq sd i i i i cossin sincos (1) The PMSM mathematical model in the d-q frame [2] is Frsq sd sssr srss sq sd Ni i pLRLN LNpLR v v 0 (2) where sdv is the stator voltage in d-axis, sqv is the stator voltage in q axis, isd is the stator current in d-axis, isq is the stator current in q-axis. The equation (5) can be stated as sd sqsqrsdssd sd L iLNiRV i dt d (3) sq Frsqssdrsdsq sq L NiRLNiV i dt d (4) The electric torque of PMSM is sqsdsqsdsqFe iiLLiT (5) where Lsd is the stator inductance in d-axis and Lsq is the stator inductance in q-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002408_978-3-319-99522-9_9-Figure9.17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002408_978-3-319-99522-9_9-Figure9.17-1.png", + "caption": "Fig. 9.17 The 6-6 Stewart parallel manipulator", + "texts": [ + "36), is equal to the degrees of freedom associated with all the moving links m \u00bc 36 minus the total number of independent constraint relations l \u00bc 30 imposed by the joints. The mechanism of the manipulator consists of six chains, including six variable lengths with identical topology, all connecting the fixed base to the moving platform. We suppose a circular moving platform symbolically represented by six pairs of spherical joints A4, B4, C4, D4, E4, F4 and a fixed base represented by another six universal joints located at the points A1;B1, C1, D1, E1, F1 (Fig. 9.17). Since the parallel manipulator is an assemblage of links and joints, this can be symbolised in a more abstract form known as equivalent graph representation, using the associated graph to represent the topology of the mechanism (Fig. 9.18). In the kinematical graph representation, with five independent loops, the links are denoted by vertices, the prismatic joints by thick edges, the revolute and spherical joints by thin edges and, finally, the fixed link 0 by two small concentric circles. In what follows, we consider that the moving platform is initially located at a central configuration, where the moving platform is not rotated with respect to the fixed base and the mass centre G is at an elevation OG \u00bc h above the centre of the fixed base" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.58-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.58-1.png", + "caption": "Fig. 11.58 Cornering coefficient and aligning torque coefficient. Reproduced from Ref. [51] with the permission of SAE", + "texts": [ + " Tires have been changed to have a wider tread and lower aspect ratio to improve maneuverability, because the nonlinearity of the cornering force with respect to the load is less for tires with a wider tread and lower aspect ratio. 11.9.2 TPC Specification System General Motors [51] proposed the TPC specification system in 1974. The TPC have parameters of the cornering coefficient (f-function), aligning torque coefficient, load sensitivity (h-function) and load transfer sensitivity (g-function) for evaluation of the cornering properties of a tire. The cornering coefficient (f-function) is defined as the side force produced at a slip angle of 1\u00b0 and 100% of the 24-psi rated load, divided by the rated load Fz (Fig. 11.58a). This tire parameter is the most influential tire parameter affecting the linear directional control performance of the tire/vehicle system. Similarly, the aligning torque coefficient is defined as the aligning torque at a slip angle of 1\u00b0 and 100% of the 24-psi rated load, divided by the rated load Fz (Fig. 11.58b). The amount of aligning torque generated by the tire, in conjunction with the desired amount of vehicle front-suspension aligning torque compliance, controls the directional behavior of the vehicle. In addition, the aligning torque generated is important in determining the force fed back through the steering wheel to the driver during any vehicle maneuvering. The load sensitivity (h-function) is a measure of howmuch a tire is able to increase the side force produced at a slip angle of 1\u00b0 as the load increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001105_biorob.2008.4762900-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001105_biorob.2008.4762900-Figure2-1.png", + "caption": "Fig. 2. Free-body diagram of the human-like mechanism and the resultant reaction loads without GRFs.", + "texts": [ + " Resultant Reaction Loads In a simulation environment where the GRFs cannot be measured directly, the ZMP should be calculated from the applied loads. This subsection describes an efficient formulation of the resultant reaction loads, which includes the gravity, the externally applied loads, and the inertia. The Lagrange\u2019s equation, which is derived for the whole mechanism including the global DOFs, yields the globalDOF generalized torques as well as the mechanism\u2019s joint actuator torques without the GRFs application (Fig. 2). Then the resultant reaction loads applied at the body reference point can be derived from the global-DOF generalized torques. The ZMP and the GRFs derived from the resultant reaction loads are applied to the mechanism in the current configuration to naturally remove the global-DOF generalized torques, which ensures a physical consistency. Since the origins of the last three local frames for the global rotations are collocated with the body reference point, the 6 \u00d7 6 Jacobian matrix J of the body reference point with respect to the global-DOF generalized coordinates is a block diagonal matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002493_s11277-018-5949-1-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002493_s11277-018-5949-1-Figure4-1.png", + "caption": "Fig. 4 The robot moves to create a new regional coordinate system, as indicated by landmarks", + "texts": [ + " If the recognized landmark did not be in a previous memory, a new node is established and a regional coordinate system is set up. If it is node , the regional coordinate system 1 3 is represented by \u03a9 = \u27e8 v, \u27e9 . The representation of the regional coordinate system creates the spatial landmark relations as follows. 4. In this case, the robot makes spatial landmark relations about the recognized landmarks around it. If no other landmarks are observed, then the semantic mapping is complete with the current node. Figure\u00a03 shows the results, which can be expressed as 5. As indicated by (d) in Fig.\u00a04, the robot approaches to the new landmark to estimate distance and angle. If the landmark did not find on a previous memory, a new node is established, as showed by the regional coordinate system for (e) in Fig.\u00a05. 6. The robot creates spatial landmark relations of the landmarks at position (e) in Fig.\u00a04. If no more landmarks appear, the robot has completed its semantic-metric map with the current node. { \u0398 , S } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v )}} 1\u0393 = { \u0398 , S , } a,b\u22081O{ \u0398 , S , } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v ) , ( 1 , 1 )}} { \u0398 , S , } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v ) , ( 1 , 1 )}} 2\u03a9 = \u27e82v, \u27e9 \ufffd \u0398 , S \ufffd = \ufffd \ufffd 2 , 2v \ufffd , \ufffd \ufffd 2 , 2v \ufffd , \ufffd 2 , 2v \ufffd\ufffd\ufffd 1 3 7. Spatial node relations are made from the previous node to the current one as follows: 1,2s v = (||| 1X \u2212 2X ||| ) , 1,2s v = ( \u2220 ( 1X \u2212 2X )) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002907_2018-32-0052-Figure17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002907_2018-32-0052-Figure17-1.png", + "caption": "FIGURE 17 Structural model of tire", + "texts": [ + " The engine and other heavy loads were assumed to be\u00a0rigid bodies and placed at their center of gravity positions and coupled by springs. Rotary joints were used to couple the steering system to the frame head pipe, the vehicle frame to the swing arm pivot, and front and rear axles. Regarding the coupling between the vehicle and the rider, the same method as that used in the simulation of the excitation experiment on the test bench was used. Figure 16 shows an external view of the model. \u00a9 2018 SAE International and \u00a9 2018 SAE Japan. All Rights Reserved. The FEM tire model provided by DYNA as shown in Figure 17 was used. This model consists of several representative elements. The material properties of each element were calculated by the numerical formula of the reference [8] based on the structure of the actual tire. Table 5 and Figure 18 show the characteristics of the tire itself obtained from the FEM tire model. Camber stiffness shows a value about 15% higher FIGURE 15 Comparison of test and simulation TABLE 4 Specifications of simulation model Wheelbase (mm) 1575 Caster angle (deg) 27.5 Weight of motorcycle (kg) 230 Weight of steering assy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003792_b978-0-12-812667-7.00031-8-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003792_b978-0-12-812667-7.00031-8-Figure3-1.png", + "caption": "FIG. 3", + "texts": [ + " One of the biggest challenges with the development of implantable biosensors is biofouling [60, 61]. Nearly all implantable biosensors that work well for in vitro tests cannot withstand the rigors of the in vivo environment [62]. The body\u2019s natural response is to fibrotically confine the implanted device and prevent it from interacting with the surrounding tissue [61, 63\u201365]. The implanted sensor becomes fouled with a film of proteins and cells on the order of 10\u2013100\u03bcm thickness within a short period of time following implantation (Fig. 3A). This cellular encapsulation forms a mass transfer barrier for analyte (glucose, lactate, etc.) diffusion to the sensing element, thereby degrading the in vivo 8354 CNT-BASED BIOSENSOR sensor performance and long-term stability [55]. Some types of biosensors can also lose their function due to long-term exposure to body fluids. Many approaches have been explored to tackle the biofouling problem [66\u201369]. Currently, there is no solution that is effective for different types of sensors [60, 61]. The active biosensor may provide a reliable way to protect the sensor from biofouling and contamination by body fluids. An active biosensor can change its configuration from closed to open to take a measurement, as shown in Fig. 3B and C. The active sensor is also computer controlled, and depending on the value of the measurement, the sensor can take additional measurements, cycle to clean the electrode, or close until the next measurement is desired. The active biosensor can be a simple bioimpedance sensor as shown in Fig. 3A. The bioimpedance sensor can be used to measure the impedance of prostate tissue as an exploratory approach for cancer detection. Tissue electric impedance is a function of its architecture. The complex impedance of each kind of tissue depends on frequency in a characteristic manner. The impedance data of various tissues in different states (normal, altered by ischemia, or cancerous) show that the characterizing differences occur at frequencies below 500kHz and down to a few kilohertz [70]. Bioimpedance has been used to differentiate normal and cancer tissues in a variety of organs including the breast [71], cervix [72], skin [73], prostate [74, 75], and bladder [76]", + " With special front-end components installed, the biosensor can perform special purpose sensing in surgery [77\u201389]. The biosensor is 2\u20135mm in diameter depending on the application and can provide a new tool for use in medical problems. The sensor can also be used to apply current to kill cancer cells [87]. Overall, the active biosensor provides a new platform for biosensors that extends the life of the biosensor by preventing biofouling and contamination [81\u201392] and miniaturizes electronic devices. The CNT fiber-based biosensor operation is illustrated in Fig. 3. Fig. 3A shows how sensors become fouled. Fig. 3B is a biosensor with the electrode enclosed and protected. Fig. 3B shows the electrode extended to make a measurement. CNT electric wire can be used to form the solenoid to extend the sensor electrode. Other components can be fabricated with microelectromechanical system (MEMS) technology. The biosensor can be powered electrically through CNT electric fiber running across the skin or by a pair of coils at the inside and outside of the skin. In another application, the biosensor can be used on a probe or needle where the nanotube wire does not cross the skin. Electrical Communication to the Biosensor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000640_icma.2009.5246170-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000640_icma.2009.5246170-Figure2-1.png", + "caption": "Fig. 2. Initial position of the machine , zero shift middle legs position (left), front shift middle legs position 20 sm (right).", + "texts": [ + " Number of the rows of matrix Do are m + 6 x 4 = 48 and more of number degrees of freedom n - k = 30 . The cause of the static indefinite is three dimension contact force at every finger contact point. To solve one indefinite proposed include add big mass at contact point, linked with finger via ball joint. Redundant drivers may be transform similarly by include two or three parts kinematics chains. The first investigated regime of motion is the motion with a shift legs middle position to front as show in Fig. 2. At Fig. 3 is presented program vertical reaction at one of the finger for some shift values. Note, from Fig.3, that the sign of reaction at finger for zero shift is positive. At the case of shift legs middle position the reactions may have negative sign - Fig. 3. The negative sign of the vertical reaction is denote physical unrealizability at control movement. So at this case the control movement may be unstable. At Fig. 4 shows the animation of the control movement for zero legs shift and shift 10 sm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.3-1.png", + "caption": "Figure 1.3 The regimes of line contact lubrication, adapted from Winer WO, Cheng HS. Film thickness, contact stress and surface temperature. In: Peterson M, Winer WO, editors. Wear control handbook. American Society of Mechanical Engineers; 1980, p. 81 141. The solid arrow represents an increase in u from 0.1 to 10 m/s at pH5 1.0 GPa. The broken arrow is u5 1 m=s and pH increasing from 0.6 to 1.6 GPa.", + "texts": [ + "2 is the case of piezoviscous-elastic lubrication. Johnson constructed a chart in gV and gE coordinates to delineate the four regimes of line contact lubrication. Hooke [4] adjusted the boundaries of Johnson\u2019s regime chart to fill in some missing portions. The boundaries in such charts are generally constructed by finding the relationships between gV and gE that make the theoretical solutions for film thickness equal for the two regimes at the boundary. Winer and Cheng [5] have shown the boundaries as \u201cfuzzy\u201d in Fig. 1.3. The regime chart in Fig. 1.3 indicates that the type of regime changes from isoviscous to piezoviscous as gV increases, and changes from rigid to elastic as gE increases. The arrows in the figure illustrate an example of a typical line contact. Here, R5 0:02 m, E05 211 GPa, \u03b1 5 15 GPa21, and \u03bco5 0:08 Pa s. The solid arrow in Fig. 1.3 represents a rolling velocity increasing from 0.1 to 10 m/s with pH 5 1:0 GPa and the theoretical film thickness grows from 0.081 to 2.04 \u00b5m as the velocity increases. The broken arrow indicates pressure increasing from 0.6 to 1.6 GPa while fixing u5 1 m=s and here the film diminishes from 0.46 to 0.36 \u00b5m for this pressure increase. These conditions are typical for steel lubricated by a mineral oil and illustrate the usual behavior for the piezoviscous-elastic regime or full EHL; the film thickness is much more sensitive to velocity (and viscosity and pressure viscosity coefficient) than it is to pressure (or load or elastic modulus). Johnson [3] also contributed dimensionless viscosity and elasticity parameters for point contacts. These are written as gV 5 \u00f0\u00f02=3\u00de\u03c03\u00de3\u03b1 p 9 H R2 E06u2\u03bc0 2 (1.6) gE 5 \u00f0\u00f02=3\u00de\u03c03\u00de\u00f08=3\u00dep 8 H R2 E06u2\u03bc0 2 (1.7) Hamrock [6] has constructed regime charts for elliptical point contacts that function in the same way as Fig. 1.3 except that the dimensionless viscosity and elasticity parameters are defined by Eqs. (1.6 and 1.7). These charts span ellipticity ratios from 0.5 to 6. Ellipticity ratio is the contact patch length in the direction normal to the motion divided by the length in the direction of motion. Large ellipticity ratio results in a thicker film by mitigating the flow of liquid out of the sides of the contact. Now it must be pointed out that the formulations, as they are given here and to follow, do not characterize the contact load by the distributed force,W, normal to the surfaces as is conventionally done, but by the maximum pressure in dry contact, pH , instead" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000714_robio.2009.4913083-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000714_robio.2009.4913083-Figure2-1.png", + "caption": "Fig. 2 Concept of support system for anteversion posture", + "texts": [ + " has been developing so called \u201cSmart Suit\u201c for the same purpose though [10], assist power is conveyed in the direction from the shoulder to the hip and it is not directly transferred to supporting torque for anteversion posture. As the result, only 13% support at maximum is realized. In this paper, Section 2 describes the concept in order to realize anteversion support, and Section 3 explains configuration of the anteversion posture support system for real use. Section 4 illustrates estimation results about how effective the support system is by the experiment to keep anteversion posture with load. As shown in Fig.2 (a), to lift from the back is one of convincing ways to support anteversion posture. A wearer can put one's full weight on the belt in this manner. Because anteversion motion is a kind of rotation one in terms of the upper and lower body around the waist, mechanical structure to realize it is developed as shown in Fig.2 (b). In this figure, inner black circle is fixed on the body and outer circle connected to the back frame rotates around it. The McKibben artificial muscle is chosen as an actuator for its light weight, flexibility, and large output. Because the McKibben artificial muscle shortens in straightaway, it is mounted at the back frame and by using a wire, it connects to outer circle by using a outer for a wire. Note that the outer is fixed to back frame as shown in Fig.2 (c) so that linear motion of the McKibben artificial muscle is converted to rotation motion. By controlling the elongation of the McKibben artificial muscle, anteversion motion is supported. How to fix inner black circle on the body in Fig.2 (b) is very crucial for this mechanism, since if it rotates, support torque for anteversion is not obtained. Fig.2 (c) describes the solution, i.e., use of waist belt and thigh belt. Waist belt is used as a normal belt for pants. When the McKibben artificial muscle is constricted, inner black circle is forced to rotate with respect to back frame. Waist frame directly connected to the inner black circle is applied and by using thigh belt through connecting belt, waist frame is fixed to the body in order to avoid rotation. In this mechanism, the front part of thigh is burdened by thigh belt during anteversion posture support", + ", t t b Mg T TbMg cos sin cossin (3) The load for horizontal and vertical direction of the front part of thigh ( xT and yT respectively) are as follows; sinsin cos sin sin Mg cb Mg TT t t tx (4) sincos cos sin cos Mg bb Mg TT t t ty (5) As a result, one leg receives 2xT and 2yT . Let\u2019s here check the concrete value. Because of sin function, 2xT and 2yT will be maximum when 2 . From the average information of adult body build, M , , b , c denote as 30kg, 0.3m, 0.2m, 0.26m respectively. Thus maximum values will be ][5.22][5.2202 ][3.17][6.1692 kgfNT kgfNT y x . (6) These values seem to be heavy load. Whereas as we described in Fig.2 (c), thigh belt has an area, i.e., about 0.1m x 0.2m corresponding to the front part of thigh. From this point of view, loads of horizontal direction and vertical one will be 86.5gf and 112.5g per 1cm2. It seems to be not so hard. In face, weraers feel it is bearable. As mentioned above, the McKibben artificial muscle is applied as an actuator. The McKibben artificial muscle was developed in the 1950s and 1960s for artificial limb research [11]. It is small, lightweight, simple, soft, flexible, and has no stiction [12]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000021_imece2009-11165-Figure21-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000021_imece2009-11165-Figure21-1.png", + "caption": "FIG. 21: AUTONOMOUS FRICTION ROBOT AND CLAMPED", + "texts": [ + " In between the rear drive wheels, a test wheel is situated, whose normal load is controlled up to 150 N as well as the slippage, which can be chosen between 0 and 100% in both rotational directions. At this solid rubber wheel the normal force and tangential force are measured and the coefficient of friction is calculated. This can be used to obtain coefficient of friction characteristics in dependence on the slippage, the relative velocity and normal load. The width of the test wheel is 18 mm, and the outer diameter is 80 mm for the unloaded state. These dimensions agree with a test wheel of a Grosch-Abrader. Fig. 21 shows a picture of this mobile test rig. In the lower right corner a picture of a so-called Grosch-wheel is shown. The material of the test wheel is Styrene Butadiene Rubber (SBR) with 60 phr Silica as filler. To perform thermography investigations on the temperature distribution directly within the rolling contact of the test wheel, a friction surface is available, which has been prepared with a germanium window. For the first experiments, the relative velocity of the friction robot is set to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002153_978-3-319-96181-1_12-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002153_978-3-319-96181-1_12-Figure5-1.png", + "caption": "Fig. 5. Eccentricity defect", + "texts": [ + " Michalec (1966) considered the case of a single eccentric gear and showed that the transmission kinematic error was a deterministic perturbation of frequency fd the frequency of defect which is equal to the frequency of rotation fr of the pinion. The amplitude is proportional to its eccentricity. E22 is the distance between the axis of rotation and the axis of inertia of the wheel and expressed by: e22\u00f0t\u00de \u00bc e22 sin X22t k22\u00f0 \u00de \u00f04\u00de Where X22 \u00bc 2pfd (fd: frequency of defect) e22 and k22 are respectively the amplitude of eccentricity and the phase of eccen- tricity. They are shown in Fig. 5. The eccentricity defect affects the potential energies and kinetic energy. In fact, this defect affects the tooth deflections. So, there is an additional potential energy which is modelled by an additional force: Fp ecc \u00bc Km t\u00f0 \u00dee12 t\u00f0 \u00de 0 0 0 rb21 rb22 0 0 0 sin a\u00f0 \u00de cos a\u00f0 \u00de sin a\u00f0 \u00de cos a\u00f0 \u00def g \u00f05\u00de The additional kinetic energy is also modelled as an additional force: FK ecc \u00bc m22e22X 2 22 0 0 0 0 0 0 0 0 cos\u00f0X22t k22\u00de sin\u00f0X22t k22\u00de 0 0f g \u00f06\u00de The equation of motion is obtained using Lagrange formalism: M\u00bd \u20acq\u00fe Cm\u00bd \u00fe Cs\u00bd \u00f0 \u00de _q\u00fe K t\u00f0 \u00de\u00bd \u00fe Ks\u00bd \u00f0 \u00deq \u00bc F t\u00f0 \u00de \u00f07\u00de [M] is the global mass matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001731_012021-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001731_012021-Figure1-1.png", + "caption": "Figure 1. Systematic diagram of squeeze film bearing with porous circular plate", + "texts": [ + " According to the Stokes micro-continuum theory of couple stress fluids together with the variation of viscosity with pressure, the combined effects of piezo-viscous dependency and nonNewtonian couple stresses in parallel circular pours plate squeeze-film characteristics are presented in this paper. The squeeze film pressure is obtained. Comparing with the iso-viscous Newtonian-lubricant case, the squeeze film characteristic of parallel circular pours plates are presented and discussed through the variation of the viscosity pressure parameter and the non-Newtonian couple-stress parameter. In Fig. 1, a systematic diagram of squeeze film lubrication between porous circular plates with normal velocity dt dh and approaching each other is shown. 3 1234567890 \u2018\u2019\u201c\u201d The lubricant is taken to be a Stokes couple stress fluid. Under the assumption, fluid inertia, body forces and body couples are negligible,in the following analysis we further assume that the viscosity varies with pressure only, the basic equations governing the lubricant velocity and pressure reduce to 0 1 z w ur rr (2) r p z u z u 4 4 2 2 (3) 0 z p (4) The above equations are solved under the boundary conditions 0 2 2 z u u and 0 w at 0 z (5) 0 2 2 z u u and t h w at hz (6) where u w x z h is the viscosity, p is the pressure, denotes the material constant accountable for non- Newtonian couple stress fluid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001483_978-3-319-72730-1_17-Figure17.9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001483_978-3-319-72730-1_17-Figure17.9-1.png", + "caption": "Fig. 17.9 Slip equation: sign of \u03b3", + "texts": [ + "5 The other conditions for field orientation stipulate that these two components of the stator current must be controlled independently. Note, however, that (R\u2032/s) \u00b7 I = j\u03c9L \u2032 r \u00b7 I m , or the two components must always satisfy the slip equation (here in static form) R\u2032 \u00b7 I = (\u03c9s) \u00b7 j L \u2032 r \u00b7 I m (17.4) 4In the \u201cL\u201d circuit of Part1, we referred the total leakage to the rotor. 5Is it now safe to conclude that there is always field orientation in an induction machine? The two components are therefore narrowly related to the slip frequency, while their vectorial sum is equal to the stator current (see also Fig. 17.9). Ifwe succeed in controlling both components of the stator current, then both torque and flux are controlled. However, in contrast to a synchronous machine, the flux of an induction machine cannot be measured directly, as has already been mentioned (e.g. the flux rotates at slip frequency with respect to the rotor and thus the rotor position is not representative of the flux position). To realise field orientation for an inductionmachine, there are twomethods: direct field orientation and indirect field orientation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002336_siela.2018.8447093-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002336_siela.2018.8447093-Figure8-1.png", + "caption": "Fig. 8 The distribution of the magnetic induction in the generator with two rotors and one stator.", + "texts": [ + " 6 shows the distribution of the magnetic induction for the generator with one stator and two rotors, while fig. 7 presents the integral magnetic induction per m3 Bc and the volume of the stator winding Vc. From these values the integral magnetic induction for the volume of the stator winding is calculated - 0,133 ,cB T\u2032\u2032 = . By means of the expressions from (3) to (9) the r.m.s value of the induced e.m.f. in one winding of the generator with one rotor and two stators is defined. . 30 2 c mc p \u0415 NA B n \u03c0\u2032\u2032 \u2032\u2032 \u2032\u2032= . (10) Fig. 8 gives the distribution of the magnetic induction in the generator with two rotors and one stator, while fig. 9 illustrates the integral magnetic induction per m3 - Bc,, as well as the volume of the stator winding Vc., by means of which the integral magnetic induction for the volume of the stator winding is calculated - 0, 202 ,cB T\u2032\u2032\u2032 = . Again by means of the expressions from (3) to (10) the r.m.s. value of the induced e.mf. in one winding of the generator with two rotors and one stator is defined " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002529_012004-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002529_012004-Figure6-1.png", + "caption": "Figure 6. Contours of three most severe stress components in the local coordinate system normalized by the respective strength for uniaxial tension along the global x-axis", + "texts": [ + " Stresses in the global coordinate system are denoted as , while stresses in the local coordinate system are denoted by a prime, . To quantify the severity of stresses within the tows, components of stress in the local coordinate system are normalized by a nominal strength and denoted by a hat, . Ideally, the strengths would be based on a series of experiments or microscale damage analyses. However, for this paper, the strengths will be assumed to be the same as the tows tested in Ref. [6], which have similar elastic properties to those used herein. 3.2.1 Tension Along X-Axis. Figure 6 shows the three most severe normalized stresses within the clipped analysis region for an applied volume average strain along the global x-axis, , of 1%. The stresses shown are in the local coordinate system, and each stress component is normalized by the respective strength, which provides a measure of the severity of each stress component. Overall, Figure 6 shows that wefts and binders experience the most severe stresses. The binders experience severe and components of stress, as shown in Figure 6b and Figure 6c respectively. The severe occurs where the binders travel through the thickness of the textile model, as shown in Figure 6b. To show the locations of stress concentrations more clearly, Figure 7 shows within an - slice of one of the binders. In this region, the local z-axis almost aligns with the global x-axis, which is the direction of the load, refer to Figure 7b. Additionally, there are no wefts in the region where the binders travel through the thickness of the textile. Consequently, the binders take on much of the load via tension along the local z-axis. The peak stresses occur where wefts are nearby to transfer the load to the binders", + " This component of stress will likely cause transverse matrix cracking within the binders. 10 1234567890\u2018\u2019\u201c\u201d IOP Conf. Series: Materials Science and Engineering 406 (2 18) 012004 doi:10.1088/1757-899X/406/1/012004 11 1234567890\u2018\u2019\u201c\u201d IOP Conf. Series: Materials Science and Engineering 406 (2 18) 012004 doi:10.1088/1757-899X/406/1/012004 In addition to the severe transverse tension, the binders experience a severe shear stress, , where they begin and finish traveling through the thickness of the textile, as shown in Figure 6c. The shear stress develops to maintain equilibrium as the binders shift between carrying a significant amount of load at the top and bottom of the textile, where the fibers of the binders are aligned with the load direction, to carrying much less load as they travel through the thickness of the textile. Though the binders experience severe stresses, the wefts experience the most severe stress, namely , within the textile, as shown in Figure 6a. For the wefts, the local y-axis is closely aligned with the global xaxis, which is the direction of the applied load for this configuration, so it is expected that the transverse tension in the wefts will be severe. However, the severity of the transverse tension varies dramatically depending on the location within the wefts. Figure 8 shows the local and global stress components of interest for an x-z cross-section centered on a binder tow, which is illustrated in Figure 7a. Figure 8a shows the contours in the wefts and binder in the x-z cross-section" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002489_s13198-018-0747-4-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002489_s13198-018-0747-4-Figure1-1.png", + "caption": "Fig. 1 The location of the equilibrium elbow", + "texts": [ + " Section 5 provides a numerical example. Section 6 concludes the paper. The equilibrium elbow is an important component of transport the force in the tracked vehicle. It combines the vehicle\u2019s body with the equilibrium elbow shaft and the road wheel shaft which can transport torsion loads. When the tracked vehicle passing the obstacle the equilibrium elbow will be torsion and make torsion shaft deformation, the deformation will absorb the energy of loading shock and it is an important process in vehicle transport system (Fig. 1). The 3D model and structure of equilibrium elbow as shown in Figs. 2 and 3. From the Figs. 2 and 3 we can see there is a hole in the middle of equilibrium elbow which radius is marked with R2, it combines the equilibrium elbow and the shock absorber to buffer the vibration of vehicle body. In Fig. 4 three relative position of the equilibrium elbow and vehicle body has been showed in order to explain the function of the equilibrium elbow clearly. From the Fig. 4 we can see three position of the equilibrium elbow compare with the vehicle body and the location of the limiter is the ultimate position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002287_s00170-018-2422-y-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002287_s00170-018-2422-y-Figure7-1.png", + "caption": "Fig. 7 Surface subdivision. (a) Vertices at the first axial depth, (b) Subdivision of front surface", + "texts": [ + " Each bicubic surface patch is constructed by four neighboring compensated cutter contact points along two adjacent toolpaths. The purpose of using subdivision is to represent the cutter contact surface by interpolating through a grid of necessary cutter contact points. The subdivision process starts from the boundary of each side of the post-bending workpiece material, and is implemented at each increment of the axial depth successively. With a given axial depth of cut, the first bicubic patch is generated by vertices on the outmost boundary. As illustrated in Fig. 7a, the first patch on the post-bendingworkpiece model is constructed with vertices P0-1, P1-1, P0-4, and P1-4, and is denoted as Patch (P0-1, P1-1, P0-4, P1-4). The corresponding compensated surface patch is constructed with compensated vertices P\u20320-1, P\u20321-1, P\u20320-4, and P\u20321-4, and is denoted as Patch (P\u20320-1, P\u20321-1, P\u20320-4, P\u20321-4) (Fig. 7b). The subdivision is first conducted on the transitional points between flat and curve surfaces, then on the mid-points along longitudinal edges. Therefore, the necessity of further subdivision is first examined at the transitional points P0-2, P1-2, P0-3, and P1-3 in Fig. 7a. If the deviations from compensated transitional points P\u20320-2, P\u20321-2, P\u20320-3, and P\u20321-3 towards the compensated bicubic patch defined by P\u20320-1, P\u20321-1, P\u20320-4, and P\u20321-4 are larger than a predefined tolerance value, the first patch Patch (P\u20320-1, P\u20321-1, P\u20320-4, P\u20321-4) will be subdivided into three patches bordered at the compensated transitional points P\u20320-2, P\u20321-2, P\u20320-3, and P\u20321-3 (Fig. 7b). These three subdivided patches are Patch (P\u20320-1, P\u203211, P\u20320-2, P\u20321-2), Patch (P\u20320-2, P\u20321-2, P\u20320-3, P\u20321-3), and Patch (P\u20320-3, P\u20321-3, P\u20320-4, P\u20321-4). If the deviations from points P0-2, P1-2, P0-3, and P1-3 towards Patch (P\u20320-1, P\u20321-1, P\u20320-4, P\u20321-4) are within the predefined tolerance value, no further subdivision is needed. The subdivision process then proceeds to the next level of axial depth of cut. When surface patch Patch (P\u20320-1, P\u20321-1, P\u20320-4, P\u20321-4) is subdivided into Patch (P\u20320-1, P\u20321-1, P\u20320-2, P\u20321-2), Patch (P\u20320-2, P\u20321-2, P\u20320-3, P\u20321-3), and Patch (P\u20320-3, P\u20321-3, P\u20320-4, P\u20321-4), further subdivision is then checked at mid-points on the longitudinal boundary of each sub-patch as illustrated in Fig. 7b. The subdivision continues until no further division is needed for all sub patches. The subdivision process then proceeds to the next level of axial depth until the final axial depth is reached. The flowchart in Fig. 8 summarizes the subdivision process. A bending machine and a CNC milling machine are sequentially applied on workpieces to implement the bendingmachining hybrid process. Three design models made of aluminum 6061 with a thickness of 0.125 in. were prototyped with this hybrid process approach" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003692_s1068799818040104-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003692_s1068799818040104-Figure5-1.png", + "caption": "Fig. 5. The roller finite element and the adjacent beam elements modeling the rings.", + "texts": [ + " The artificial low-order magnitude stiffness xxK is added to each node of the outer ring to eliminate its rigid body motions. The distributed load 0( ) = cosxq q b\u03d5 \u03d5 is applied along the elements of the inner ring with the node coordinates > 0x such that RUSSIAN AERONAUTICS Vol. 61 No. 4 2018 573 2 0 2 = cos .xF q b d \u03c0 \u03c0 \u2212 \u03d5 \u03d5\u222b (19) The problem is two dimensional in the plane XY , so the degrees of freedom = = = 0z x yu \u03b8 \u03b8 vanish at all the nodes. At the location of the j th roller a specially developed 2-node finite element shown in Fig. 5, describing the behaviour of the rolling element compressed between the two rings, is placed. The element has two external degrees of freedom (one per each node\u2014 iw and ew ) and one internal degree of freedom rw . A local coordinate system r , t , z (see Fig, 1) is defined for each element with the corresponding unitary vectors, which form the local rotation matrix .T T T r x t y z zR = e e + e e + e e (20) The deflections along the axis r (see Fig. 5) are computed in the local coordinate system via the dot product = , = ,e H r i G rw w \u2212u e u e (21) where Gu and Hu are the displacement vectors of nodes G and H in the global coordinates. The interference of the rolling element surface into the rings can be calculated as follows: = ; = . 4 4i r i e r e g g w w w w\u03b4 \u2212 \u2212 \u2212 \u03b4 \u2212 + \u2212 (22) The unknown radial displacement rw is determined while searching for the roller equilibrium in accordance with (8) using the Newton\u2013Raphson method. The local stiffness coefficients are required in order to successfully treat this equation: 1= " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.1-1.png", + "caption": "Fig. 18.1 Two rigid bodies in relative motion maintaining contact", + "texts": [ + " Before considering sliding or rolling of deformable bodies, it is helpful to explore the most general classes of motion permitted between two rigid bodies if they are constrained to remain in contact. Only the relative kinematics is of importance to the contact problem,1 so it is convenient to define a frame of reference which moves so as to ensure that the instantaneous contact point is always located at the origin and the common normal between the contacting bodies is always vertical. It follows that the only non-zero velocities of particles adjacent to the contact point in the two-dimensional example of Fig. 18.1 are the tangential velocities V1, V2. If V1 =V2, we can define the sliding velocity VS = V2 \u2212 V1. (18.1) If there is a tangential force (e.g. a frictional force) Q transmitted between the contacting bodies, the power dissipated in friction will be |QVS| and the force must oppose the sliding motion. 1We are not concerned with dynamic effects for the moment, so there is no need to [e.g.] restrict attention to an inertial frame. \u00a9 Springer International Publishing AG 2018 J.R. Barber, Contact Mechanics, Solid Mechanics and Its Applications 250, https://doi.org/10.1007/978-3-319-70939-0_18 433 If V1=V2\u2261VR = 0, the sliding velocity VS =0 andwe have a state of pure rolling. The tangential force Q will then transmit power between the bodies at a rate |QVR|, but there will be no frictional losses. Since the contact point has no vertical component of velocity in Fig. 18.1, the instantaneous centres of rotation O1, O2 of the two bodies must lie on the common normal, and in fact they must be located at the centre of curvature of the surface immediately adjacent to the contact point, since if this condition were not satisfied, material points just upstream of the contact point would move to locations above or below the stationary contact point. Notice that the bodies do not have to be cylinders\u2014only the local region at the contact must be cylindrical [see Fig. 18.2] and the instantaneous centres O1, O2 will generally move along the common normal as sliding and/or rolling proceeds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002507_icicct.2018.8473139-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002507_icicct.2018.8473139-Figure2-1.png", + "caption": "Fig. 2. Pendulum motion with motor torque", + "texts": [], + "surrounding_texts": [ + "Board 2.0 for NI ELVIS. The RIP system is a multivariable mode with highly nonlinear behavior. It is used in the design of highly complex applications like automatic aircraft landing system, and humanoid robot stabilization. In this work, the RIP model is mathematically derived using motion equation of EulerLagrange. For controlling the system, PP (pole placement) and LQR(Linear Quadratic Regulator) controllers are designed and implemented in MATLAB for the balancing of the pendulum arm in up-right position. A comparative study was made for analyzing the performance of the controllers in terms of settling time and overshoot range, etc. Based on designed parameters using MATLAB simulation, the control ofreal time Rotary inverted pendulum arm is implemented in LabVIEW platform interfaced with NI ELVIS. From the experimental results of MATLAB simulation and LabVIEWinterfacing, the LQR controller was better than the PP controller in controlling the pendulum arm in upright position. Keywords-Rotary Inverted Pendulum, QNET 2.0 Rotary Pendulum Board for NI ELVIS, PP, LQR, MATLAB, LabVIEW. I.INTRODUCTION Inverted Pendulum is abench mark model and classically problem to control the nonlinear system dynamics in control engineering application. There are two types of inverted pendulum model available as the models to define the nonlinear dynamics in translational and rotational system. In this paper Rotary inverted pendulum is used to define the system with highly nonlinear and open loop multi variable unstable system. The application of RIP model in controlling the altitude of space booster rocket, automatic aircraft landing system, and satellite[1][2]. RIP invented by Furuta, is also known as Furuta's pendulum. It has two moving components with amotor driven arm and a free moving pendulum arm. The problem of balancing the pendulum at uprightpositionhas been divided into two major controls i.e., swing upcontrol and balance control. The swing up control is used tobring the pendulum from its stable position to itsvertical upright position and balance control stabilizes thependulum at its upright position. The system model was derived using Euler-Lagrange method as it is a multivariable system[2][3]. Since the dynamics of inverted pendulum systems are inherently nonlinear, the equations ofmotion are linearized about the operating point and is defined within which the constant gaincontroller results in local asymptotic stability. QNET 2.0 is a rotary inverted pendulum interfaced in LABVIEW interfaced using NI-ELVIS data acquisition card. The motor driven arm is the base of the system which rotates in the horizontal plane and a pendulum attached with that arm which is free to rotate in the vertical plane[4]. The pendulum angles are measured using two encoderswith a resolution of 0.176deg/count. In this paper, pole placement and LQRcontrollers are designed for the RIP model. The pole placement controller places the system poles in desired position for obtain a desired response. The LQR method is another powerful technique to obtain a feedback gain for balancing the pendulum in minimum cost of function.The LQR technique generates controllers with guaranteed closed-loop stability and robustness property[1].The paper is organized as follows. The mathematical modelof the rotary inverted pendulum is presented in the Section II.The controllerdesigns for LQR and Pole placement is discussed in section III. Simulationresults of RIP model with controller design in MATLAB simulation and 978-1-5386-1974-2/18/$31.00 \u00a92018 IEEE 1452 LABVIEW real time implementation are presented in Section IV with conclusion inSection V. II. STANDARD MODEL The modeling of rotary inverted pendulum is a challenging one because it is a multivariable system. Normally it is a nonlinear system. So the modeling is done by using the motion equation of Euler-Lagrange. The following assumptions are taken in modeling of the system: 1. The system starts in the state of equilibrium that means the initial conditions are assumed to be zero. 2. The pendulum deflection is restricted within few degrees in up-right position to satisfy the linearity property. 3. Pendulum can be subjected to small disturbance input. 4. The pendulum is displaced at an angle of \u03b1 in y-axis while the arm angle of \u03b8 considered in x-axis. TABLE I. PARAMETERS OF RIP MODEL DESCRIPTION Symbol Description L1 =Lp/2 Length of the Pendulum (Lp) from its Center of Mass(m). Mr Mass of pendulum rotary Arm (kg). Mp Mass of Pendulum (kg). Lr Pendulum rotary Arm Length(m). Pendulum Arm Deflection (radian). \u03b1 Pendulum Deflection (radian). Jp Inertia of Pendulum about its center of mass(kgm 2 ). Jr Inertia of Pendulum rotary Arm about its center of mass(kgm 2 ). \u03c4 Servo motor torque(Nm). Kt Motor torque constant (Nm/A). Km Back emf constant (V/rad/sec). Rm Armature resistance(ohm). Vm Motor input voltage(Volt). Vx Velocity of Pendulum about its Center of mass in the x-direction. Vy Velocity of Pendulum about its Center of mass in the y-direction. The pendulum experiences the velocity due to the center mass of the pendulum in two directions(Vpen) Vpen= -L1cos\u03b1*(\u03b1\u0307)*x-L1sin \u03b1*(\u03b1\u0307)*y (1) The pendulum rotary arm also experiences the velocity only in horizontal direction(Varm) as in [1] Varm=Lr\u03b8\u0307 *x (2) From the equation (1) and (2) the velocity (V) experienced by the pendulum is based on the resultant of two axis. Vx= Lr\u03b8\u0307- L1cos\u03b1 * (\u03b1\u0307) (3) Vy= -L1sin \u03b1 * (\u03b1\u0307) (4) The dynamic equation of the system is obtained by the variation of velocity in the pendulum which can be derived using Euler-Lagrange motion equation[3]. By Euler-Lagrange equation of motion Qi ( \u0307 ) \u0307 (5) Where, q = The angular position. \u0307 =The angular velocity. Qi= External force. L = Lagrangian. W= loss of energy. Lagrangian is a function of angular position and angular velocity, which is represented as L=Ttotal- Ptotal (5a) Ttotal =Total kinetic energy of the RIP Ptotal = Total potential energy of the RIP. Total potential energy in the system is only due to the gravity in pendulum, so potential Energy is given as Ptotal=Parm+Ppend (5b) As the arm of the pendulum is already balanced, Parm=0 Ptotal=MpgL1cos\u03b1 (6) Total kinetic energy (KE)ofthe system can be obtained by the sum of kinetic energies due to the pendulum arm, velocities of pendulumdue to center of mass, and rotational motion of pendulum, Ttotal =KEpend.arm+KEVx+ KEVy + KEpend (7) Ttotal=(1/2)*Jr \u0307 2 +(1/2)*Mp[Lr \u0307-L1cos\u03b1*( \u0307)] 2 +(1/2)*Mp[-L1sin\u03b1*( \u0307)] 2 +(1/2)*Jp \u0307 2 After getting the kinetic energy and potential energy, the lagrangian is given as L=(1/2)*Jr \u0307 2 + (1/2)*MpLr 2 \u03072 - MpLr \u0307 L1cos\u03b1*( \u0307) + (1/2)*Mp L1 2 cos 2 \u03b1*( \u0307) 2 + (1/2) MpL1 2 sin 2 \u03b1*( \u0307) 2 + (1/2)Jp \u0307 2 - Mpg L1cos\u03b1 (8) Normally,Jp=(1/12)*MpLp 2 andLp=2*L1 Hence, Jp=(1/3)MpL1 2 (9) Substituting equation (9) in (8) L=(1/2)*Jr \u0307 2 + (1/2)*MpLr 2 \u03072 - MpLr \u0307 L1cos\u03b1*( \u0307) + (1/2)*Mp L1 2 cos 2 \u03b1*( \u0307) 2 + (1/2) MpL1 2 sin 2 \u03b1*( \u0307) 2 + (1/6)*MpL1 2 \u03072 - Mpg L1cos\u03b1 (10) 978-1-5386-1974-2/18/$31.00 \u00a92018 IEEE 1453 Parameters Values Mr 0.095 kg Mp 0.024 kg Lr 0.085 m Lp 0.129 m L1= Lp/2 0.0645 m Kt 0.042 Nm/A Km 0.042V/(rad/sec) Rm 8.4 ohm Jp 0.0001 Kgm 2 Jr 5.7*10 -5 Kgm 2 After substituting the parameters in the expression, A=0.0002304, B=0.00013158, C=0.000133128 D=0.01518, E=1.335E-8, F=0.3528 and G=0.00021 Computed state equation of the RIP model is [ \u03b8\u0307 \u03b1\u0307 \u03b8\u0308 \u03b1\u0308 ] =[ ] [ \u03b8 \u03b1 \u03b8\u0307 \u03b1\u0307 ]+ [ ]Vm (19) The output matrix of the RIP model is [ ] =[ ] [ \u03b8 \u03b1 \u03b8\u0307 \u03b1\u0307 ] (20) III. CONTROLLER DESIGN AND IMPLEMENTATION" + ] + }, + { + "image_filename": "designv11_92_0000371_pime_conf_1963_178_384_02-Figure16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000371_pime_conf_1963_178_384_02-Figure16-1.png", + "caption": "Fig. 16..3. Particle of oil considered in converging film", + "texts": [ + " (2) White-metal wiping criterion BEARING HYDRODYNAMIC THEORY indicates that the load carrying capacity of a journal bearing under ideal conditions tends to infinity as the minimum oil film thickness approaches zero. In practice, however, the minimum effective hydrodynamic oil film thickness in the journal bearing would have a limiting value, and hence the maximum load capacity would be finite. The limiting value of film thickness depends on several factors. Ocvirk and DuBois (I)? suggest that when the highest asperities of the two mating surfaces make initial contact a load carrying film exists. The minimum effective film thickness at this point being the sum of the \u2018predominant peak roughness\u2019 (Fig. 16.1) of the The M S . of this paper was first received a t the Institution on 12th September 1963 and in its revised form, as accepted by the Council for publication, on 14th February 1964. * Research and Development Organization, The Glacier Metal Company Limited, Alperton, Wembley, Middlesex. t References are given in Appendix 16.11. Proc Instn Mech Engrs 1963-64 two surfaces. This occurs where friction due to metalto-metal contact starts to increase, and so for safe running the minimum allowable film thickness should be greater than this figure", + " Ocvirk and DuBois (I) have examined experimental results and determined the Sommerfeld number at the points of minimum coefficient of friction for various bearings with different surface finishes. They assumed that these points were where metal-to-metal contact began. Calculations were then made for oil film thickness at these minimum friction conditions using the short bearing theory, and then compared with the \u2018predominant peak roughnesses\u2019. The predominant peak roughness was based on work by Tarasov (3) who examined the surface roughnesses produced by different machining processes and found two orders of roughness, a predominant peak roughness and the peak to valley value shown in Fig. 16.1. He compared the predominant peak roughness values with the profilometer roughness (R.M.S.) values and this ratio is referred to as the Tarasov factor which varied with different machining techniques. Ocvirk found that the sum of the \u2018predominant peak roughnesses\u2019 (h, in Fig. 16.1) using Tarasov factors gave comparable results with those calculated from bearing theory. It therefore seems reasonable that the minimum allowable film thickness should be at least greater than the sum of the \u2018predominant peak roughness\u2019 of the two mating surfaces. The British system of measuring surface finish is the centre-line average (c.1.a.) system and, therefore, Tarasov factors do not directly apply. However, Rubert (4) has produced a relation between R,,, and c.1.a. values after studying tests on 200 samples (where R,,, is the peak to valley value of surface finish)", + " The value of the peak to valley surface finish R,,, is seen to be dependent on both the c.1.a. value and the type of machine process involved and both these factors will probably depend on the size of the bearing. For instance, a 2-in diameter bearing may conceivably be diamond bored, and the journal cylindrically ground, both having a surface finish of the order of 8 to 16 p-in c.1.a. On the JOURNAL INANT GHNESS BEARING Rmax,, Peak to valley surface finish (p in) of journal. R,,,, Peak to valley surface finish (p-in) of bearing. Fig. 16.1. Idealized oil film thickness h, where initial Conversion factor ($) =m 62.4 1 metal contact is made 1 1728 lb (force) in sec ( 1bcS X - = 4 ~ 1 0 - ~ at NATIONAL SUN YAT-SEN UNIV on August 23, 2014pcp.sagepub.comDownloaded from MINIMUM ALLOWABLE OIL FILM THICKNESS IN STEADILY LOADED JOURNAL BEARINGS I63 1 2 3 4 5 7 8 9 other hand, a 20-in diameter bearing would be finished bored and the journal finished turned with surface finishes of the order of 32 to 63 p i n c.1.a. A survey was carried out to obtain some idea of representative finishes for various journal diameters. Gladman (5) associated values of surface finish with various machine operations and P.E.R.A. (6) have shown similar relations together with machining tolerances. This coupled with experience of machining processes and tolerances used in bearing manufacture gave some guidance to representative surface finishes as indicated in Fig. 16.2. The dotted lines show the \u2018spread\u2019 of results obtained and when considering the manufacture of bearings one would tend to keep to the lower level of this range. Typical surface finishes associated with various journal diameters are shown in Fig. 16.2, and this can be used as a general rough guide. Experience indicates that in many instances surface finishes very much better than those shown have been produced in bearings, and these should therefore operate with a greater margin of safety. 8 10.3 2.8 5.7 7.7 6.0 6.5 9.8 Proc Instn Mech Engrs 1963-64 The purpose of associating surface finish with machining processes and bearing sizes is to obtain a very simple quick reference guide, and owing to its simplicity it must be very approximate in nature, but should still give the right trends for guidance in estimating an allowable minimum film thickness. It is assumed that the \u2018predominant peak roughness\u2019 is half the peak to valley value R,,,, and also the surface finish for the journal and bearing are assumed to be of the same order. This results in a minimum film thickness of h, equal to R,,, at metal-tometal contact conditions. Using Rubert\u2019s Rmax/c.l.a. relation from Table 16.1, a graph showing the general trend of R,,, with various journal diameters is shown in Fig. 16.2. Neglecting shaft deflection and maximum oil film temperature considerations, a general guide for the minimum allowable film thickness based on surface finish would be twice (say) the R,,, value shown in Fig. 16.2. The surface finish criterion is based on current machining techniques, and is a factor to some extent under the control of the designer in specifying the surface finish he requires. Specific cases of surface finish which differ from those of Fig. 16.2 can be used together with the RmaX/c.l.a. relation in Table 16.1 to obtain some idea of minimum allowable film thickness. Experimental verification Davis (7) carried out some tests on 2-in diameter bearings having different clearances. These bearings were \u2018run in\u2019, but before doing so, the values of ZN/P (units of cP, rev/min per lb/in2) giving minimum coefficient of friction were recorded. It is these values of ZN/P which are considered, giving a corresponding theoretical hydrodynamic minimum film thickness which is compared with the initial surface finish measurements", + " WHITE-METAL WIPING CRITERION At high surface speeds and high loads, the bearing metal may become plastic and \u2018wipe\u2019 under local high-temperature conditions in the oil film. As a rough guide, it is assumed that wiping occurs at the minimum film thickness position. This is not strictly true, but the associated high oil film pressures causing plastic flow will not be far removed from the minimum film thickness position, especially when the eccentricity ratio is very high and the attitude angle approaches zero. It is, therefore, necessary to determine the maximum oil film temperature at the minimum film thickness position (4 = T in Fig. 16.3). Consider what happens to a small particle of oil travelling through a converging clearance space over a 180\u201d arc (i.e. from 4 = 0 to C$ = T) , the particle of oil being at the centre of the bearing width and, therefore, having no axial Proc Instn Mecli Engn 196344 component of flow (assuming a parallel clearance space axially). The problem is approached by considering the frictional heat created in the particle being transferred by the circumferential oil flow. The following assumptions are made : (1) Heat by conduction is neglected in this theoretical analysis; a conduction factor, however, may be used if substantiated by experimental results. (2) The temperature-viscosity characteristics of the oil can be represented by an equation of the form 0 = A/vq where A and q are constants. It so happens that most lubricating oils approximate to this law at temperatures greater than 70\u00b0F as shown in Fig. 16.4 provided the units are (temperature) \u201cF and (kinematic viscosity) cS. (3) This analysis assumes that the velocity profile of the oil at any position circumferentially is linear. This is not strictly true, but by making this assumption a relatively simple approximate solution is obtained rather than a more exact unmanageable solution. From the oil specification the rate of change of temperature with viscosity dO/dv is determined for a given value of Vol178 Pr 3N at NATIONAL SUN YAT-SEN UNIV on August 23, 2014pcp.sagepub.comDownloaded from IMINIMUM ALLOWABLE OIL FILM THICKNESS IN STEADILY LOADED JOURNAL BEARINGS 165 viscosity. This combined with the heat equation for the oil particle considered is then integrated around the bearing arc to determine the temperature distribution. Several typical oils are shown in Fig. 16.4 with values of q varying from 0.325 for the heavier turbine oil A to 0.404 for the lighter turbine oil E. The rate of change of temperature with change in viscosity from equation 0 = A/vq is given by: (16.1) Consider the particle of oil (Fig. 16.3) travelling around the converging oil film at bearing mid-width position. The rate of heat generated in this particle is given by: U2r d#J dz Q P = V hJ (Btulsec) . (16.2) The circumferential rate of heat transfer by the oil is given by : Qt = vbPch do U 2 = - hpC, dz dB (Btulsec) . (16.3) Equating (16.2) and (16.3), substituting for dB from equation (16. l) and putting h = C( l + E cos 4) gives : The integration of the above equation and the evaluation of constants are shown in Appendix 16.1", + "5) will be cS, a unit commonly used by bearing designers. It should be noted that the ratio B,/Bo is not dimensionless in the general sense, since 8 must be in units of O F . The maximum oil film temperature (assumed to be at the minimum film thickness position) will depend on all the factors shown inequation (16.5). To reduce thenumber of factors involved, the clearance ratio can be associated with journal speed. Normally at high-speed operation bearings are designed with higher clearances and a general guide for average clearances is shown in Fig. 16.5. The numerical relation between clearance ratio and journal speed N (revlmin) taken from Fig. 16.5 is C NO.2 - X 103 = - . . . (16.6) r 3.34 Substituting this in equation (16.5) together with J, - 9 E and 4, and transposing gives the following minimum film thickness ratio for maximum temperature conditions at position #J = n: (3 where x is a convenient substitution such that x [T JCh 8,r (:)I (dimensionless) 3.5 x 1010 6, (1+4)/4 Ch 80 n2 I(&) -1j [&]Nu, - x, therefore, depends on the temperature of the oil at minimum and maximum film thickness positions and on the oil characteristics. A graph showing the minimum film thickness to diameter ratio against x for various journal speeds from 100 to 50000 rev/min is shown in Fig. 16.6. Brown and Newman (9) carried out some tests on a plain bearing of 6 in diameter operating at 6000 revlmin. The usual form of failure was found to be overheating of the white metal and it was noted that the critical temperature in most instances was about 340-350'F. The white-metal bearing may wipe at different temperatures depending on the developed pressure in the oil film in the high-temperature region of the bearing. The higher the temperature becomes the smaller the pressure required to give permanent plastic flow of the material. For bearings running at very large eccentricity ratios, the peak pressures in the oil film will be large, and therefore the critical temperature may well be less than 340-350\u00b0F. The following data is used to obtain a relation between the minimum oil film thickness and journal diameters, and speeds for a maximum oil film temperature 8, of 300'F: Assumed temperature Oo at maximum film position Oil specification: Turbine oil C. Viscosity characteristics as Fig. 16.4. Viscosity vo at 150\u00b0F equals 21 cS. Oil constant q equals 0.364. Assumed specific heat C, equals 0.45 Btu/lb degF. = 150\u00b0F. Val I78 Pr 3N Prac lnstn Mech Engrs 1963-64 at NATIONAL SUN YAT-SEN UNIV on August 23, 2014pcp.sagepub.comDownloaded from 166 F. A. MARTIN by Burke and Neale (2), it can be assumed that the calculated minimum oil film thickness from their procedure will be at the bearing mid-width position for a misaligned journal. The journal deflection over halfthe bearing width must, therefore, be taken into account. The deflection ( T ~ , ~ over half the bearing width, for a SPEED Fig. 16.6. Film thickness ratio h,/d for various values of x For the above conditions the value ofx equals 3.71 x lolo, and applying this to Fig. 16.6, a direct relation between the minimum film thickness to journal diameter ratio (h,/d) and journal speed ( N ) is obtained, which is used in constructing the film thickness guidance diagram, Fig. 16.7. A composite guidance graph for minimum idealized oil film thicknesses at conditions of metal-to-metal contact between journal and bearing, and also for the minimum film thickness at maximum temperature conditions of 300\"F, is shown in Fig. 16.7. This graph shows that at high surface velocities of 300 ft/sec, the white-metal wiping criterion gives a limiting film thickness of about four times that obtained on a basis of surface finish. This shows the importance of considering the maximum local temperature in the oil film and also shows that improvement in surface finish does not directly help in the higher speed white-metal wiping region. For the purpose of calculating the load carrying capacity of a bearing using such design methods as shown Proc Instn Mech Engrs 1963-64 Vol 178 Pt 3N at NATIONAL SUN YAT-SEN UNIV on August 23, 2014pcp.sagepub.comDownloaded from h4INIMUiM ALLOWABLE OIL FILM THICKNESS IN STEADILY LOADED JOURNAL BEARINGS 167 simple rotor system as shown in Fig. 16.7 may be obtained from the equation where K6 is a factor depending on the position of the load between the two bearings and also on the ratio of bearing width to the distance between the two bearing centrelines. Values of Kd are shown in Fig. 16.8. As a guide to a safe value of minimum oil film thickness it is suggested that Allowable hmin = 2hf+6,,, , where h, is obtained from Fig. 16.7 and 6,,, may be obtained from Fig. 16.8, if such a rotor system is applicable. It must be emphasized that Fig. 16.7 gives only a rough guide to bearing designers in fixing the design limit of journal position in a bearing. It is hoped that its usefulness will lie in its simplicity, but by virtue of its simplicity it is necessarily only of an approximate nature, due to the assumptions made in achieving such a diagram. The author would therefore appreciate any discussion or communication giving experimental results so that the diagram may be checked against them. The author would like to thank The Glacier Metal Co" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.28-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.28-1.png", + "caption": "FIGURE 16.28 (a) Laser-arc welding system with both laser beam and arc on the same side of the workpiece. (b) Laser-arc welding system with the laser beam and arc on opposite sides of the workpiece. (From Duley, W. W., 1983, Laser Processing and Analysis of Materials. By permission of Springer Science and Business Media.)", + "texts": [ + "3\u20132 times than that obtained using the 5 kW laser alone. Similarly, higher penetration is obtained with the laser augmented gas metal arc welding (LAGMAW) process than with gas metal arc welding (GMAW) alone. The increase in penetration associated with LAGMAW is more significant at low arc welding currents (Fig. 16.27). A similar behavior is observed at lower speeds. The reduction in penetration at higher currents could be due to absorption of the laser beam by the increased plasma formation. Possible arrangements of the setup are illustrated in Fig. 16.28. It is desirable to have the distance between the laser beam and the electrode as close as possible. In the case of GTAW, no significant increases in penetration is observed when the separation is less than 3 mm and the laser power is 2 kW with a current of 300 A, and a welding speed of 1 m/min. At greater distances, however, the penetration reduces with increasing separation distance. Even though the electric arc can be located on either side of the workpiece, it has been found to be more effective on the side of the workpiece opposite the laser (Fig. 16.28b). In addition to increased penetration, another advantage of the augmented welding process is the elimination of undercut and humping that can occur at high speeds during GTAW alone. One disadvantage of arc-augmented welds is that they tend to have a larger heat-affected zone size than a weld made using a laser of the same total power. When both the laser and arc are on the same side of the workpiece, the bead appearance is found to be better when the laser beam leads the arc in LAGMAW. This is because the suppression gas flow does not affect the molten pool created by the arc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000822_icinfa.2009.5204980-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000822_icinfa.2009.5204980-Figure2-1.png", + "caption": "Fig. 2. Description of speed vector on a 3D inertial system of coordinates.", + "texts": [ + " The rest of the paper is organized as follows: In section II, mechanical analysis of human movement describes the speed deployment of an action performed under an inverse dynamics model. Section III presents six linear and non-linear classifier representations. Experimental results are shown in section IV whereas section V points out conclusions and future work. 978-1-4244-3608-8/09/$25.00 \u00a9 2009 IEEE. 528 II. MECHANICAL ANALYSIS OF HUMAN MOVEMENT The mechanical movement of a human limb can be rigorous in its description of how an actor performs some physical activity within an inertial system of coordinates. Fig. 2 illustrates this notion where under the expression of a physical activity, the major speed feature vector u\u0307 is shown to be produced from the trajectory C through the distance \u0394r over the time \u0394t. Similar to the theoretical vector analysis of [6], in Fig. 2, during the punching activity we observe that the actor\u2019s fist is displaced through the curve C which is the trajectory of a punch. The position vector of point p at time t is r = r(t), while the position vector of point q at time t +\u0394t is r+\u0394r = r(t +\u0394t). The speed of the fist at p is given by Equation 1: v = dr dt = lim \u0394t\u21920 \u0394r \u0394t = lim \u0394t\u21920 r(t + \u0394t)\u2212 r(t) \u0394t (1) which is a vector of p in C. If r = r(t) = x(t)i+y(t)j+z(t)k are the right vector components in coordinates x, y, and z, given by Equation 1 while the speed is given by Equation 3: v = dr dt = dx dt i+ dy dt j+ dz dt k (2) u = |v| = \u2223\u2223\u2223\u2223dr dt \u2223\u2223\u2223\u2223 = \u221a( dx dt )2 + ( dy dt )2 + ( dz dt )2 = ds dt = u\u0307 (3) The speed vector u\u0307 is regarded as the major mechanical attribute of this analysis because it is the base feature used by all the other feature categories shown on Table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000107_1475921708094791-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000107_1475921708094791-Figure1-1.png", + "caption": "Figure 1 A gear grid cross-section showing a tooth region with reduced rigidity.", + "texts": [ + " But, having conducted the grid independence study using the present methodology for steady state solutions earlier [8], the results presented here should be sufficiently accurate in the present context. Also, as shown in Kaul [8], a small measure of Coriolis effect may be present even at 6000 rpm, during the transient state. The second simulation is the repeat of the first simulation except that rigidity of one of the gear teeth is decreased in a certain fashion over a at MCGILL UNIVERSITY LIBRARY on February 12, 2015shm.sagepub.comDownloaded from portion as shown in Figure 1. For example, the rigidity of the particular tooth is made spatially variable in both the circumferential and radial directions extending all the way from the inner radius to the outer radius. The variation is parabolic in the circumferential direction and exponential in the radial direction. The value of rigidity at the mid section at the outer radius is half the nominal value, and it increases parabolically outward circumferentially to the nominal value. Radially, the rigidity decreases from this value from the outer radius inward to the inner radius by a factor of (1.0 \u2013 exp( 0.25 i)), where i\u00bc 1 at the inner radius and i 1 at the outer radius. This may be representative of a damage state that would yield distinctly different vibration signature from the gear than the normal. The results from these two simulations are compared side by side in terms of the time evolution of radial, tangential and shear stresses at selected radial locations on the radial line running through the middle of the two selected gear teeth, tooth #1 and tooth #10, as shown in Figure 1. Since the damaged tooth is made less rigid, its vibrations are distinctly different from those of the normal tooth, but these vibrations attenuate more rapidly than those from the normal case. Both the teeth vibrate with the same frequency, which is the natural frequency of the system, about 3.0KHz. The radial and tangential directions in the present case correspond to the two orthogonal generalized curvilinear coordinate directions, as shown in Figure 1. A comparison of tangential stress evolution at three radial locations along the radial line running through the middle of the two teeth is shown in Figures 2\u20134. According to theoretical steady state rotating annulus results [15,16], the tangential stress is maximum at the inner radius and decreases monotonically toward the outer radius, which is also the behavior exhibited by the tangential stress distribution along the radius in the simulation here. Since the results correspond to an approximate steady state, the results for teeth 1 and 10 are almost identical for the normal (undamaged) gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002392_icca.2018.8444363-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002392_icca.2018.8444363-Figure1-1.png", + "caption": "Fig. 1. The Prototype Variable-Pitch Quadrotor", + "texts": [ + " This paper aims to develop a proper lift controller with optimized power consumption for a prototype variable-pitch quadrotor, based on which the control quality and the endurance performance can be improved. The rest part is organized as follows. Sec. II is the problem overview, in which the prototype quadrotor, the modelling and optimal power curve analysis are presented. Sec. III discusses the nonlinear adaptive RPM control law and its implementation, followed by the overall lift controller in Sec. IV. Sec. V presents the test results and Sec. VI draws a conclusion. The prototype variable-pitch quadrotor drone is shown in Fig. 1. Different from conventional fixed-pitch quadrotors, it\u2019s equipped with four variable-pitch propellers, four electric motors and the torque-transmitting mechanisms. For detailed parameters, one can refer to our previous work [25]. The variable-pitch propellers driven by the electric motors 978-1-5386-6089-8/18/$31.00 \u00a92018 IEEE 1046 through the torque-transmitting mechanisms can provide sufficient force to lift the vehicle weighed from 8 kg for no load to 15 kg for full load. As for the four variable-pitch propellers of the prototype quadrotor, they have the same characteristics and can be controlled independently" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000068_j.mechmachtheory.2008.05.003-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000068_j.mechmachtheory.2008.05.003-Figure4-1.png", + "caption": "Fig. 4. Tangency of inflection circles in the two conjugate positions.", + "texts": [ + " Thus requiring the inflection circles of the conjugate positions associated with a double point of centrode to be tangent at the instant center provides a reasonable general approach towards the study of self tangency for the pf-curve of any planar linkage. Disproving the possibility for the satisfaction of such conditions seems a promising approach towards our objective. The extensions of the links 2 and 4, in the first position intersect the distinct inflection circles at (JA, JB) and at \u00f0J0A; J 0 B\u00de in the second conjugate position, as illustrated in Fig. 4. Since the extensions of the links 2 and 4 remain unchanged in the two positions, considering homothety about I, we can write: j \u00bc IJB IJA \u00bc IJ0B IJ0A \u00f021\u00de On the other hand, having the extensions of links 2 and 4 unchanged together with Eq. (21) satisfied, necessitates the homothety of the two conjugate positions associated with the double point. These conditions provide necessary requirements for self tangency. Let us proceed applying the Euler\u2013Savary equation [7] for the inflection circle of the FBL at the both positions: IJB \u00bc IOB\u00f0IOB BOB\u00de BOB ; IJA \u00bc IOA\u00f0IOA AOA\u00de AOA \u00f022-I; II\u00de Applying the Sine law to the triangle OAIOB as previously shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000505_adem.200800032-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000505_adem.200800032-Figure1-1.png", + "caption": "Fig. 1. Geometry of smooth (left) and notched (right) micro specimens (all values are given in lm). Areas a and b see text below.", + "texts": [ + "[5,6] Thereto a polyoxymethylen (POM)-model of the specimen was made using a micro milled mould. 15 \u201cPOM-specimens\u201d were put on a cast-tree and embedded in gypsum. After dewaxing and sintering, the mould was heated and evacuated. The preheating temperature Tm of the mould was 1000 \u00b0C. In order to compare notched and smooth specimen of aluminum bronze, each third specimen on the cast tree had no notches to avoid a systematic influence of the position of specimen at the cast-tree. The specimen geometry is shown in Figure 1. The cross section intended to have nominal values of 260 lm in width and 130 lm in depth. Two different notch geometries were designed with radii of 12 and 50 lm. Due to this geometry the notch factors kt of 2 and 3.5, respectively, should be achieved. Two batches (A and B) of tensile specimens were produced. Each batch contained five smooth specimens, five with radii of 12 and five with radii of 50 lm. C O M M U N IC A TIO N S ADVANCED ENGINEERING MATERIALS 2008, 10, No. 6 \u00a9 2008 WILEY-VCH Verlag GmbH & Co", + " To optimize the contrast of the specimen\u2019s surface, which was necessary to have a good contrast for the gray scale correlation analysis of the strains, the specimens were etched for 5 sec with ferric nitride. During the tensile test the surface structure changes the reflection property. So it was not possible to evaluate all specimens using the grey-scale correlation method. The notch factor gives the relation between the stress in the notch tip and the nominal stress. To evaluate this notch factor the local strain was calculated using the smallest possible area (\u201cb\u201d in Fig. 1 which had a width between 20 and 40 lm) around the tip of the notch. Due to the experimental setup it was not possible to look at the notch root directly. Nominal strain was calculated within the area \u201ca\u201d marked in Figure 1 and adapted with respect to the different cross sections. Results and Discussion The values of width and depth scatters in a range of about 7 % resulting from the production process. The characterization of the micro specimen regarding the surface topography showed the values for width in and out of the notch as well as the depth and roughness parameters, which are listed in Table 2. Values of roughness parameters are the same for notched and smooth specimens. Specimens of Batch B had a slightly larger cross section and greater roughness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002798_detc2018-85998-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002798_detc2018-85998-Figure1-1.png", + "caption": "FIGURE 1. THE PROPOSED PLANAR COMPLIANT TENSEGRITY MECHANISM (COTM) COMPRISES OF TWO TRIANGULAR RIGID BODIES CONNECTED BY A RIGID ROD (GREEN) AND TWO SPRING MEMBERS (RED)", + "texts": [ + " For a ZFL spring, the unstretched (free) length of a linear spring is zero and a general ZFL spring is not an existing product [21]. This greatly simplifies the analysis of the system, however this assumption is not valid for actual systems. The paper presents detailed analysis for an example planar CoTM where ZFL assumption is gradually relaxed 1) both springs have ZFL, 2) one spring has ZFL and 3) no springs have ZFLs. The proposed mechanism consists of two triangular rigid bodies connected by a rigid rod and two springs members as shown in Fig. 1 where the relative distance between points 1,4 is L3. The spring free lengths and spring constants are denoted by L0i,ki where i = 1,2. The objective is to find all of its stable equilibrium positions, as defined by \u03b31 and \u03b32, for a given set of parameters. Let the coordinate system B be fixed on the bottom rigid body with origin at point 1, x-axis between points 1, 2 and z-axis out of the plane of paper. Similarly, the coordinate system T is fixed on the top rigid body with origin at point 4, x-axis between points 4, 6 and z-axis out of the plane of paper. The points are BP2 = [p2x,0,0] T , BP3 = [p3x, p3y,0] T (1) T P5 = [p5x,0,0] T , T P6 = [ p6x, p6y,0 ]T (2) The transformation matrix T BT between coordinate system T and B is written as T BT = c2 \u2212s2 0 L3c1 s2 c2 0 L3s1 0 0 1 0 0 0 0 1 (3) where ci,si are cos\u03b3i,sin\u03b3i corresponding to angles \u03b31,\u03b32 shown in Fig. 1. Hence, for i = 4,5 [BPi 1 ] = T BT [T Pi 1 ] (4) The superscript will be dropped for remaining section of the paper as all the calculations will be performed in the B coordinate system. The static equilibrium equations for force and torque balance yield F = f1 ( P2\u21926 d1 ) + f2 ( P3\u21925 d2 ) + f3 ( P1\u21924 L3 ) = 0 F = f1 d1 (P1\u21924\u00d7 P2\u21926)+ f2 d2 (P1\u21924\u00d7 P3\u21925) = 0 (5) \u03c4 = P1\u21922\u00d7 ( f1 \u00b7 P2\u21926 d1 ) + P1\u21923\u00d7 ( f2 \u00b7 P3\u21925 d2 ) = 0 (6) where f3 is the unknown force along the rigid bar, fi = ki(di\u2212 L0i) are the forces in spring elements and di are length of the spring elements \u2200i = 1,2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001529_s12008-018-0471-y-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001529_s12008-018-0471-y-Figure8-1.png", + "caption": "Fig. 8 The displacement cloud map. a The displacement of the bearing. b The displacement of the liner", + "texts": [ + " The mesh size of load plate and mandrel is 1 mm. The total number of nodes is 383447 and the elements are 93832. The analysis element is solid 185, and this element is defined by 8 nodes having three degrees of freedom at each node. The boundary conditions are as follows: the both ends of mandrel are fixed completely; the outer ring surface of the load plate is set to full constraint; the radial load of 44KN is applied to the load plate. The finite element model is shown in Fig. 7: The simulation results are shown in Fig. 8. The simulation results from Fig. 8 showed that the maximum displacement between the liner and inner ring occurred on the axial cross-section, this section is the intersectionplane between XOY planes. Thence, the maximum wear depth is on the same section. With the intention of simplify the calculation and the debugging of the program, the middle section where themaximumdisplacement between the inner ring and the outer ring of the self-lubricating joint bearing is taken as an object. Since the rigidity of the mandrel is large and the deformation is small, the mandrel is ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002125_978-981-13-0411-8_71-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002125_978-981-13-0411-8_71-Figure7-1.png", + "caption": "Fig. 7. FE8 test rig [17]", + "texts": [ + " The contact and wear models presented in this paper apply to contact processes without intermediate medium (dry friction). However, for a film thickness parameter k\\0:25 the solid body contact is assumed to be predominant (interfacial friction) [16]. Using this assumption, the simulation model is validated by means of experimental tests. The tests are carried out at low lubricant quantities and low relative speeds. The results are evaluated by profilometry and weighing. The experiments are run on a FE8 test rig, shown in Fig. 7. In this test rig, two test bearings (2) were mounted on a common shaft, the housing washers are fixed. The shaft was radially supported by two supporting bearings and driven by an electric motor. A static load was applied axially via a cup spring assembly (1). The design ensures that both bearings were subjected to the same load. Volumetric wear was determined by tactile measurement of the profile before and after tests. The gravimetric wear was determined by weighing the bearing components. As lubricant, a mineral oil with a viscosity m40 \u00bc 186 cSt at 40 C was used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002068_yac.2018.8406526-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002068_yac.2018.8406526-Figure1-1.png", + "caption": "Fig. 1. Linear motor inverted pendulum", + "texts": [ + "3, for any \u03b3 > 0 and any compact set \u0393 \u2282 Rq containing the origin of Rq , there exist three layer feedforward neural network functions x\u0302(W, v), u\u0302(W, v) satisfying (7). As a result, a control law of the form g(x, v) = u\u0302(W, v) + L(x\u2212 x\u0302(W, v)) (8) where L is such that all the eigenvalues the matrix (\u2202f/\u2202x)(0, 0, 0)+(\u2202f/\u2202u)(0, 0, 0)L are inside the unit circle solves the approximate discrete-time nonlinear output regulation problem of system (1). In the section, we will establish the discrete-time mathematical model of the linear motor inverted pendulum and then give the problem formulation. Consider the linear motor inverted pendulum shown by Figure 1, where x \u2208 R represents the position of the base of the pendulum in the horizontal plane, \u03b8 \u2208 (\u2212\u03c0, \u03c0) represents the angle the pendulum makes with vertical, u \u2208 R is the control force being applied to the cart, l is the length from the base of the pendulum to the center of mass, M is the mass of the cart driven by linear motor, m is the mass of the uniform pendulum, and I is the moment of inertia of the pendulum. The dynamic model can be deduced using Lagrangian dynamics and is given as follows [15]: (M +m)x\u0308+ml(\u03b8\u0308 cos \u03b8 \u2212 \u03b8\u03072 sin \u03b8) = u (I +ml2)\u03b8\u0308 \u2212mgl sin \u03b8 +mlx\u0308 cos \u03b8 = 0 (9) Let x1 = x, x2 = x\u0307, x3 = \u03b8, x4 = \u03b8\u0307, the state-space equations of the system are as follows: x\u03071 = x2 x\u03072 = 1 M +m\u2212 m2l2(cos x3)2 I+ml2 \u00b7(u+mlx2 4 sinx3 \u2212 m2gl2 cosx3 sinx3 I +ml2 ) x\u03073 = x4 x\u03074 = 1 Ml +ml(sinx3)2 + (M+m)I ml \u00b7 ( (M +m)g sinx3 \u2212u cosx3 \u2212mlx2 4 cosx3 sinx3 ) (10) By discretizing the continuous time model (10) via Euler\u2019s method with T as sampling period, the discrete-time model of the linear motor inverted pendulum is given as follows: x1(k + 1) = x1(k) + Tx2(k) x2(k + 1) = x2(k) + T M +m\u2212 m2l2(cos x3(k))2 I+ml2 \u00b7(u+mlx2 4(k) sinx3(k) \u2212m2gl2 cosx3(k) sinx3(k) I +ml2 ) x3(k + 1) = x3(k) + Tx4(k) x4(k + 1) = x4(k) + T Ml +ml(sinx3(k))2 + (M+m)I ml \u00b7 ( (M +m)g sinx3(k)\u2212 u cosx3(k) \u2212mlx2 4(k) cosx3(k) sinx3(k) ) y(k) = x1(k) (11) Remark 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003715_s11015-019-00768-0-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003715_s11015-019-00768-0-Figure2-1.png", + "caption": "Fig. 2. Dimensions of cell elements (a) and CAD-model of a BCC type specimen cellular structure (b).", + "texts": [ + " The starting materials for the SLM process used was powder of this steel (Fig. 1) with a particle size of 5\u201345 \u03bcm. The alloy chemical composition was as follows, %: C up to 0.03, Si up to 0.6, Mn up to 0.8, Cr 15.0\u201317.0, Ni 14.0\u201316.0, Mo 2.5\u20133.0; S up to 0.015, and P up to 0.02. Powder fluidity 16.1 sec (GOS\u0422 20899-98), bulk density 4.47 g/cm3 (GOS\u0422 19440-94). Shape factor 1.5 (spherical powder). For studies a cellular structure of the BCC type was used with a volume fraction of cavities of 70%. The sizes of lattice cell elements are given in Fig. 2a, and CAD model of a specimen with a cellular structure of the BCC type is given in Fig. 2b. A computer model for specimens with a cellular structure was prepared in a program for object component topological optimization developed in FGUP RFYaTs-VNII\u00c9F (Russian Federation Nuclear Center, All-Russia Scientific Research Institute of Experimental Physics). Laser power, W 100 100 100 100 100 100 Scanning rate, mm/sec 565 565 565 565 565 565 Distance between tracks, mm 0.09 \u2013 \u2013 0.09 0.09 0.09 Number of laser passes 3 1 1 2 2 2 Laser power, W 175 175 100 100 100 175 Scanning rate, mm/sec 750 750 565 565 565 750 Distance between tracks, mm 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002496_app.46884-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002496_app.46884-Figure2-1.png", + "caption": "Figure 2. (a) The model of 3D-printed test platform. (b) The 3D-printed platform embedded with test sample, permanent magnet, and microcontroller. (c) Optical and magnetic dual-responsive bending deformation was tested by using the multilayered films with 10%PVA/40%Fe3O4 and 10% PVA/5%5CB/3%Graphite electrospun fiber films. [Color figure can be viewed at wileyonlinelibrary.com]", + "texts": [ + " (In the thermal responsive deformation test, 5 wt % 5CB and 20 min electrospinning time were selected as the better parameter of the optical responsive deforming film fabrication.) The preparation of vis-light and magnetic field responsive deforming multilayered films can be seen from Figure 1(c). First, the 10 wt % PVA/Fe3O4 precursor solution contains 10, 20, 30, 40 wt % Fe3O4 were used for orderly electrospinning 20 min upon the TPU film, respectively. Then, a layer of PVA/5CB/ graphite fiber film was electrospun 20 min upon the PVA/Fe3O4 fiber film for obtaining the composite fiber film. A 3D testing platform (model shown in Figure 2(a) was designed with Creo Parametric 2.0 software) embedded with electromagnetic driven mechanical switch [Figure 2(b)] was printed by 3D printer (Dreamer, Zhejiang FlashForge Co., Ltd., Zhejiang, China). The variation of magnetic field intensity can be obtained by the movement (up and down) of permanent magnet driven by the microcontroller. The experimental photo of the optical and magnetic dualresponsive bending deforming multilayered film with TPU substrate film, 10%PVA/40%Fe3O4 electrospun fiber film, and 10% PVA/5%5CB/3% graphite electrospun fiber film can be seen from Figure 2(c). Heating process was conducted by using constant temperature heating platform (JF-956, range from 0 to 450 C). A wavelength of 405 nm laser with the output power of 98 mw/cm2 (M-33A405\u2013500-G) was chosen as spot optical source. The spot optical source was irradiated on one side of the multilayered film. Because of the influence of laser beam angle, 26 cm was selected as the better distance between the film and 46884 (2 of 7) J. APPL. POLYM. SCI. 2018, DOI: 10.1002/APP.46884 laser source for obtaining the maximum energy utilization of light field irradiation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002389_978-3-319-99010-1_8-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002389_978-3-319-99010-1_8-Figure6-1.png", + "caption": "Fig. 6. The membership function of R (k)", + "texts": [], + "surrounding_texts": [ + "The error e(k) and error change De(k) of stator current is the input of the observer, and then De\u00f0k\u00de \u00bc e\u00f0k\u00de e\u00f0k 1\u00de \u00f05\u00de e\u00f0k\u00de \u00bc I s \u00f0k\u00de Is\u00f0k\u00de \u00f06\u00de Where, k is switch state of inverter. I s \u00f0k\u00de is given stator current value, Is\u00f0k\u00de is actually measured stator current value [1] (Fig. 3). The language variable of e(k) and De(k) are defined by five fuzzy language variables {NB, NS, ZO, PS, PB}. The observer output is the stator resistance error DRs (k) [7]. DRs k\u00f0 \u00de \u00bc Rs k\u00f0 \u00de Rs k 1\u00f0 \u00de \u00f07\u00de Where, Rs (k) is the value of stator resistance this time; Rs (k \u22121) is the value of stator resistance last time [1]. The language variable DRs (k) is defined by five fuzzy language variables {NB, NS, ZO, PS, PB} [2]. 4.1 Fuzzification The regulator has two inputs, e(k), \u0394e(k) and for the fuzzyfied output R(k) as to Figs. 4, 5 and 6 [2]. The set of rules is described according to Mac Vicar with the format If- Thus, under the fuzzy rules table with two input variables according to [2] (Table 1). The choice of the inference method depends upon the static and dynamic behavior of the system to regulate, the control unit and especially on the advantages of adjustment taken into account. We have adopted the inference method Max\u2013Min because it has the advantage of being easy to implement in one hand and gives better results on the other hand [8, 12]. 4.3 Defuzzyfication The most used defuzzification methods is that of the center of attraction of balanced heights, our choice is based on the latter owing to the fact that it is easy to implement and does not require much calculation [12]." + ] + }, + { + "image_filename": "designv11_92_0003231_icsai.2018.8599363-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003231_icsai.2018.8599363-Figure2-1.png", + "caption": "Figure 2. The UAVs system\u2019s the static position on the runway of the autonomous air aircraft carrier", + "texts": [ + " When griping the pull rod under the belly of UAV, the flower gripper structure of the retractable connecting rod column firmly seizes the pull rod of the landing rod of an UAV positioning UAV in the location of the slab on the taxiing runway. When the retractable rod column of slab of air aircraft carrier shrinks to its minimum, an UAV is safely positioned on air aircraft carrier\u2019s slab. The rod is able to stretch to large passenger aircraft slide the front end of the runway. The root of the rod column (tie bar) is located in the front of the aircraft cockpit, as shown in Figure 2. In Figure 3, the model of pod and pod mounting system of the airborne that grips the pull rod on the fuselage of UAV is showed. When there are flight accidents of large aircraft on geographic region, or when large aircraft has strong oscillation and vibration, it releases a UAV on top of flight, and the UAV takes off on the taxiing runway on top. Or, an autonomous aircraft carrier (AAC) release its pod-UAV system under its wing. The UAV can accompany with flying of large passenger aircraft in the sky, and the UAV monitors and tracks large passenger aircraft flight, and passes the large aircraft flight state and routes, and starts the internal software robot (software UAV) that provides the internal state and video information of large passenger aircraft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000336_iccas.2008.4694278-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000336_iccas.2008.4694278-Figure2-1.png", + "caption": "Fig. 2 Saturation function h(\u00b7)", + "texts": [ + " The effectiveness of the design method is illustrated with a practical example of dynamic positioning(DP) system which holds the position and heading of a ship under wind disturbances by controlled thrusters. Fig.1 shows a block diagram of a system consisting of a controller dynamics K(s), a dynamics of an actuator servomechanism and a plant dynamics Gp(s). When the controller gain is determined based on LQ theory, we ignore the dynamics of actuator servomechanisms and assume that all outputs can be measured. By changing the control weight of the quadratic performance criterion stepwise under the fixed state weight, a set of LQ gain groups is decided. Fig.2 shows one of saturation functions h(\u00b7) with the magnitudes \u00b11 in the actuator servomechanism in Fig.1, by scaling the input and output of the saturator. The saturator h(\u00b7) has a linear gain \u03b2 in the input range of |q| \u2264 1/\u03b2 . If the absolute value of the input q to the saturator h(\u00b7) is constrained less than 1/\u03b1 , the sector condition h \u2208 [\u03b1 \u03b2 ] is satisfied locally. The objective of this paper is to determine the maximum absolute values 1/\u03b1 of inputs to the multiple saturation functions simultaneously as reciprocals of the minimum sector bounds \u03b1 to ensure the local absolute stability of a total system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.20-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.20-1.png", + "caption": "Fig. 18.20 Brake block with an off-centre pivot", + "texts": [ + " At the centre of the contact ellipse, the effective rolling radius [r in Fig. 18.4] is R. Use Kalker\u2019s strip theory and kinematic arguments from Sect. 18.1.1 to estimate the shear traction distribution in the contact area and the energy dissipated in frictional microslip, if the coefficient of friction is f and both bodies have elastic properties E, \u03bd. Remember that if there is no acceleration, the tangential force Q=0. 6. The non-uniform wear of the brake block of Fig. 18.14a can be reduced by moving the pivot point to the left as shown in Fig. 18.20. Determine the optimal distance it should be moved, if the unworn thickness corresponds to b = b0 and the fully worn state with uniform thickness of wear for all x corresponds to b=b1 (0= = f \u03b1cos \u03b1,sin> \u03c90 2 \u2014A dynamic analysis of a machine for vibrotransportation whose drive (a dynamic vibration exciter and a vibration system (a pan)) transfers motion to a concentrated mass by means of Coulomb friction is performed. In the first approximation of the averaging method, the dynamic characteristics and the stability of the stationary regimes of vibrotransportation under forced vibrations are studied", + " In accordance with the Routh\u2013Hurwitz criterion, these conditions have the following form: (7) After transformations, conditions (7) taking into account (3) correspond to the inequality cos\u03b1 > 0, which is satisfied due to the existence of solutions of (4). Therefore, entire resonance curve 1 (Fig. 2) is stable, including the segments that do not satisfy inequality (5), which is not related to the conditions of existence of regimes (4). 3. In the case of zero initial conditions x = 0, = 0, and = 0 (A = 0 and \u03c8 = 0) and the effect of harmonic excitation E in the system (Fig. 1), both masses first move together. The parameters of the transient and stationary regimes are determined in this case from exact formulas for the linear oscillator with mass M + m and damping b. Using as an example (Fig. 2) two main cases of the amplitude\u2013frequency characteristics of this system, curves 3 show the parameters of the stationary regimes without sliding \u03c901 = K/(M + m). Curves 2 determine in this case the onset of sliding according to (5). We now consider a transient process denoted by thin arrows with points C and D (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000118_20080706-5-kr-1001.00015-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000118_20080706-5-kr-1001.00015-Figure2-1.png", + "caption": "Fig. 2. Characteristics of reference point", + "texts": [ + " When the value of s is time varying, position of the point Pr(s(t)) also moves. In this paper, \u03b6 is defined as follow: \u03b6(q\u2032, s) = sgn (q\u2032 \u00b7 e\u20322(s)) \u2016q\u2032\u2016, (7) where \u2016 \u00b7 \u2016 shows Euclidean norm. Absolute value of \u03b6 indicates length of q\u2032. Sign of \u03b6 indicates e\u20322 element of q\u2032. i.e., if \u03b6 is positive, the plant exists in left half plane of e\u20321 axis. Former relations are satisfied for any given s(t). Moreover, we consider a special point s = sr which orthogonalize q\u2032 and e\u20321(Altafini [2002],Okajima et al [2004]). Then such Pr(sr) is named as \u201creference point\u201d (See Fig. 2). I.e., sr(t) represent the dynamics of reference point when the following equation holds for any t. q\u2032(t) \u00b7 e\u20321(sr(t)) = 0 (8) If (8) holds in initial state, we can consider sr(t) by the following state equation (See Altafini [2002]): x\u0307(t) = F (x(t), u(t)), x = [xT p , xT re] T (9) F (x, u) = fp1(xp) + fp2(xp)u h1(xp)h2(xp, u) \u2212 \u03ba\u0302r h1(xp) cos \u03b8 1 \u2212 \u03ba\u0302rz h1(xp) cos \u03b8 1 \u2212 \u03ba\u0302rz h1(xp) sin \u03b8 where xre = [\u03b8, sr, z]T . (9) includes (3),(4) and (5). We can see that z indicates signed distance from the plant and the reference point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003394_6.2019-1747-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003394_6.2019-1747-Figure6-1.png", + "caption": "Figure 6: Examples of symmetric and non-symmetric displacements implemented in the experiments.", + "texts": [ + " The tape spring hinges are each fixed at each end to 3D printed PLA plastic clamps using epoxy, and custom 3D printed mounts affix the hinge to the transducers. Custom mounting brackets are also 3D printed in PLA to mount the transducer assemblies to the testbed. The experimental procedure is as follows. The hinge configuration is incremented into the symmetric configuration and data is sampled statically. Then each non-symmetric displacement is configured and sampled statically, reseting back to the symmetric configuration between each sample. Examples of the non-symmetric configurations are displayed in Figure 6. The geometry of the fixture must be taken into account when transforming the relative position and orientation data. The fixture creates an offset of the rotation axis from a01 of 42.5 mm at both sides of the testbed. The opposite sense configuration is achieved by flipping the coupon over using a modified mounting bracket, such that the hinge frames remain in the same position relative to the motor hubs. Several data samples are collected for a given configuration and are averaged to provide one sample per configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001699_1077546318767559-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001699_1077546318767559-Figure6-1.png", + "caption": "Figure 6. Force analysis of the box.", + "texts": [ + " The complex meshing force between the gears will directly affect the whole rotor system or even its dynamic characteristics. Meantime, the gear meshing force will be transmitted to the supporting bearings, and then to the whole box which will cause the system vibration. This means the vibration of the gearbox can partly reflect the meshing force. A separated analysis method is adopted to analyze the vibration of the box in order to improve the efficiency of computation, where the box is regarded as rigid body and the vibration of the foundation is neglected. The force analysis diagram of the box is shown in Figure 6. F1R F5R and F1L F5L are the excitation forces of the left bearings and the right bearings acting on the 10 bearing seats, which are the reactive force of the supporting bearing forces and related to the displacements and velocities of 10 bearing journals of the rotor system. r1R r5R and r1L r5L are the corresponding arm of force refer to mass center of the box. Fx, Fy, Fz, Tx, Ty, and Tz are the resultant forces and resultant moments act on the box and can be calculated as F \u00bc X5 i\u00bc1 FiL \u00fe FiR\u00f0 \u00de \u00bc \u00bdFx Fy Fz T T \u00bc X5 i\u00bc1 riL FiL \u00fe riR FiR\u00f0 \u00de \u00bc \u00bdTx Ty Tz T \u00f022\u00de The dynamic equation of the box can be described as Mb \u20acub \u00fe Cb _ub \u00fe Kbub \u00bc F T \u00f023\u00de where Mb, Cb, and Kb are the mass matrix, damper matrix, and stiffness matrix, respectively, and ub is the displacement vector of the box and defined as 6-DOF, which is given by ub \u00bc xb yb zb xb yb zb T \u00f024\u00de where x,y, and z are the corresponding displacements along the three local coordinate axes of the box; xb, yb, and zb are the corresponding rotation angles around the three coordinate axes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002060_j.promfg.2018.06.045-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002060_j.promfg.2018.06.045-Figure3-1.png", + "caption": "Fig. 3: (a) Temperature distribution in the preload element, (b) thermal displacement of the preload element, (c) force-displacement characteristic curve of the preload element", + "texts": [ + " These influences can be minimized by connecting the strain gauges to a full Wheatstone bridge, similar to a measuring rosette. Alternatively, micro-Pt100 elements can be used to calculate the apparent strain from their temperature signals and compensate them. In order to be able to determine the desired states of the preload element from the sensor signals, characteristic diagrams and a control loop are used (Fig. 2 (b)). FEM simulations were carried out to evaluate the potential of the thermo-elastic preload element. The simulation results are shown in Fig. 3. The Peltier element introduces a temperature difference of up to 45 K into the preload element. Through this, preload forces of up to 3 kN can be generated. Commonly applied preload forces on spindle bearings vary between 250 N for low preloads and up to 1,5 kN for high preload. The achieved preload forces are sufficient to change the operating mode between two machining steps. The switching of the operating mode can be done simultaneously to tool changes. Thus, switching does not cause any additional non-productive time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003063_j.ifacol.2018.11.667-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003063_j.ifacol.2018.11.667-Figure3-1.png", + "caption": "Fig. 3. Configuration at instant of landing of fore foot (t = T )", + "texts": [ + " The X-position of the imaginary ZMP or foot rotation indicator (FRI) point, originally studied by Goswami (1999), indicated as point F in Fig. 2; the Xfri, is defined as the same as Eq. (3), and we calculate this value in the numerical simulations to smoothly interpolate the generated motion of the center of pressure. Note that, however, this does not give physical meaning during the SSP. Let t = T be the instant of landing of the fore foot, or the time instant of transition from the SSP to DSP. Fig. 3 illustrates the coordinate at t = T . By considering the symmetry of the generated motion in the phase space, the terminal angular positions should satisfy \u03b81(T ) =\u2212\u03b81(0) = \u03b1 2 , (9) \u03b82(T ) =\u2212\u03b82(0) = \u03b81(T ) + \u03d5 = \u03b1 2 \u2212 mgl k sin \u03b1 2 . (10) For the same reason, the terminal angular velocities should also satisfy IFAC SYROCO 2018 Budapest, Hungary, August 27-30, 2018 Fumihiko Asano / IFAC PapersOnLine 51-22 (2018) 67\u201372 69 g m2, I l k \u03b81 \u2212\u03b82 m1 \u03b1 X Z O Fig. 1. Model of passive rimless wheel with spring-loaded reaction wheel reaction wheel on the condition that the extension becomes zero where \u03b81 \u2212 \u03b82 = 0", + " The X-position of the imaginary ZMP or foot rotation indicator (FRI) point, originally studied by Goswami (1999), indicated as point F in Fig. 2; the Xfri, is defined as the same as Eq. (3), and we calculate this value in the numerical simulations to smoothly interpolate the generated motion of the center of pressure. Note that, however, this does not give physical meaning during the SSP. 2.3 Target Terminal Condition Let t = T be the instant of landing of the fore foot, or the time instant of transition from the SSP to DSP. Fig. 3 illustrates the coordinate at t = T . By considering the symmetry of the generated motion in the phase space, the terminal angular positions should satisfy \u03b81(T ) =\u2212\u03b81(0) = \u03b1 2 , (9) \u03b82(T ) =\u2212\u03b82(0) = \u03b81(T ) + \u03d5 = \u03b1 2 \u2212 mgl k sin \u03b1 2 . (10) For the same reason, the terminal angular velocities should also satisfy IFAC SYROCO 2018 Budapest, Hungary, August 27-30, 2018 \u03b8\u03071(T ) = \u03b8\u03071(0) = 0, (11) \u03b8\u03072(T ) = \u03b8\u03072(0) > 0. (12) The Xzmp at this instant should be positioned at the rear foot, that is, Xzmp(T ) = k\u03d5 mg = \u2212l sin \u03b1 2 = Xrf " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.10-1.png", + "caption": "Fig. 11.10 Tire model with a stretched string [29]", + "texts": [ + " Meanwhile, Fy does not saturate for the solid tire model with a brush model in the right figure. Furthermore, the actual tire generates an aligning torque stronger than that of the solid tire model with a simple brush model. The self-aligning torque peaks before the maximum side force is achieved. It is often considered that the loss of \u201cfeel\u201d just beyond the peak of the self-aligning torque is an indication to the driver that he or she is cornering on the limit. (1) Fundamental equations for the string model Figure 11.10 shows the string model [29-32] where a string with tension is supported by springs between the string and the wheel center and the string does not deform in the circumferential direction. Referring to Fig. 11.11, the force equilibrium of the tread element with width b and length dx yields fydx ksvdx D\u00feD\u00fe @D @x dx S1 @v @x \u00fe S1 @v @x \u00fe @2v @x2 dx \u00bc 0; \u00f011:23\u00de 3Problem 11.2. x x dx y 0 S1 S1 D fydx ksvdx b/2 b/2 dx x DD x vS 1 dx x v x vS 2 2 1 v Fig. 11.11 Force equilibrium in a tread element of the string model [29] where v is the displacement of the string in the y-direction and ks, S1, D and fy are, respectively, the lateral carcass spring rate per unit length in the circumferential direction, the tension of the string in the circumferential direction, the shear force and the side force on the contact patch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001904_012033-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001904_012033-Figure5-1.png", + "caption": "Figure 5. Some of the optimum structures obtained from the best performer UPS-EMOA.", + "texts": [ + " From Table 3, the best performer based on mean hypervolume value is UPS-EMOA while the second best and the third best are MODE and RPBILDE, respectively. For the measure of search consistency based on STD of hypervolume values, the best performer is MODE while the second best, which obtained slightly higher STD, is UPS-EMOA. Overall, the UPS-EMOA is said to be the best performer for solving the proposed walking tractor handlebar optimisation problem in this study in terms of both search convergence and search consistency. Figure 5 shows some of the optimum structures obtained from using UPS-EMOA. 7 1234567890\u2018\u2019\u201c\u201d In this work, simultaneous topology, shape and size design of a walking-tractor handlebar based on using many-objective metaheuristics (MnMHs) is presented. The design problem is posed to maximise the sum of natural frequencies of the first 5 modes, minimise structure mass and minimise construction cost subject to stress and displacement constrains with 2 loading conditions, bending and torsion. Design variables include the topology of side bar stiffeners, shape, and size of all bars" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000592_cdc.2008.4739453-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000592_cdc.2008.4739453-Figure1-1.png", + "caption": "Fig. 1. Input-output behavior for the actuator with hysteresis", + "texts": [ + " The flow set is taken to be C := { [ \u03be q ] \u2208 R n+1 : K\u03beq \u2264 u\u0304, q \u2208 Q\u2032}, (4) and the jump set is taken to be D := { [ \u03be q ] \u2208 R n+1 : K\u03beq \u2265 u\u0304, q \u2208 Q\u2032 } . (5) In the following subsection, we will describe the behavior of the system with hysteresis, which can be regarded as a special case of the aforementioned hybrid model. Consider the input-output behavior of an actuator v = u \u2212 \u01ebq, \u2200(u, q) \u2208 {(u, q) \u2208 R \u00d7Q\u2032 : uq \u2264 a}, (6) where u is the feedback control input, \u01eb > 0 is the width of the hysteresis, a > 0 is the threshold and q \u2208 Q is the logic variable. Fig. 1 shows the input-output relationship. On the other hand, when the actuator is saturated, its inputoutput behavior is shown in Fig. 2 and the representation is given as v = sat(u \u2212 \u01ebq), \u2200(u, q) \u2208 {(u, q) \u2208 R \u00d7Q\u2032 : uq \u2264 a}, (7) where \u01eb and a are defined in the same way as below (6). Assumption 1: a > 0, \u01eb > 0, \u01eb < min{1, a} and 1 \u2212 \u01eb < a \u2264 1 + \u01eb. Hence, in Fig. 2, the threshold is such that switching happens in the linear section of each curve, and after switching, the output is saturated. When a = \u01eb+1, two curves connect and the output does not jump", + " The control objective is to achieve exact tracking of constant references r = 2/3 and the closed-loop poles are all located at s = \u22121. To implement the integral control for exact tracking (see e.g., [2, page 552]), denote e = x1\u2212r and xI = \u222b edt. Then y = e + r and the system is e\u0307 x\u03072 x\u0307I = 0 1 0 0 \u22120.5 0 1 0 0 e x2 xI + 0 1 0 u. (17) The control law is given by u = \u22123e \u2212 2.5x2 \u2212 xI so as to stabilize the system (17) and achieve exact tracking for the original system. However, if such a system is subject to actuator nonlinearity as shown in Fig. 1, and Fig. 2, then the closed-loop performance will be affected. As in Section II, we describe this system with a hybrid model as follows. States: [\u03be1 := e, \u03be2 := x2, \u03be3 := xI + \u01ebq, q ] T . The flow map is given as: f(\u03be, q) = [\u03be2, \u22120.5\u03be2 + \u03c3(u\u0303), \u03be1, 0] T , where u\u0303 = u\u2212\u01ebq = [ \u22123 \u22122.5 \u22121 ] \u03be. In section IV-A, we will discuss the case \u03c3(u\u0303) = u\u0303, while in section IV-B, we will discuss the case \u03c3(u\u0303) = sat(u\u0303). The jump map is: g(\u03be, q) = [\u03be1, \u03be2, \u03be3 \u2212 2\u01ebq, \u2212q] T . The flow set and the jump set are: C := { [ \u03be q ] : [\u22123 \u2212 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003657_ceit.2018.8751889-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003657_ceit.2018.8751889-Figure1-1.png", + "caption": "Fig. 1. Longitudinal dynamics of a wheel [7]", + "texts": [ + " SYSTEM MODELLING AND VALIDATION This paper focuses on tire dynamics of a quarter car, mainly longitudinal dynamics. Longitudinal dynamics consist of translation motion of vehicle and wheel rotation. In literature, a lot of quarter car models are available [21- 23] which are easy and simple to implement in Simulink. In this section, we will present longitudinal dynamics of the quarter car model. We can use the model presented in [7], as a reference. According to Newton\u2019s 2nd law, the longitudinal and rotational dynamics of the quarter car system in fig. 1 can be represented by the following equation [21]: J\u2126 = \u0393\u2212 RFx \u2212 Cf\u03c9 Mvx = Fx \u2212 Fd \u2212 Frr (1) Different forces defined as 1) Traction Force Fx: This force comes from the tireroad interaction and mainly depends on the slip ratio and friction coefficient. It can be expressed as a function of the coefficient of friction \u03bc(\u03bb) with the ground Fx(\u03bb)=Fz\u03bc(\u03bb) (2) With \u03bb as pseudo-slip which is defined as \u03bb= R\u2126 \u2212 vx R\u2126 = 1 \u2212 vx R\u2126 (3) The relation between \u03bc and \u03bb is semi-empirical [21] \u03bc(\u03bb)=2\u03bc0 \u03bb0\u03bb \u03bb0 2+\u03bb2 (4) where \u03bb0 is pseudo slip which corresponds to the maximum value of the friction " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001140_1.5061279-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001140_1.5061279-Figure2-1.png", + "caption": "Figure 2. Relative position between laser beam, workpiece and cutting tool, (a) end-view and (b) sideview.", + "texts": [ + " o15+ o6\u2212 o45 The workpiece used was a Ti6Al4V alloy bar with a diameter of 62mm. Its optical microstructure is shown in Figure 1. It shows two phase (alpha plus beta) structure with the alpha phase grain size of 3\u03bc\u03bc7 \u00b1 and the hardness of 30Hv351\u00b1 . Page 901 of 909 s A 2.5kW Nd:YAG laser was used for heating the workpiece. The laser beam was delivered by a 15m long optical fiber and focused by an optical lens with a focal length of 200mm on the chamfer surface at the angle of with the workpiece axis (as shown in Figure 2). The distance between the focusing lens and workpiece ( ) was set so that the laser spot size covered the chamfer surface. The tool-beam distance, , was set at 10, 20 and 30mm respectively. o40=\u03b8 196mmL1 = 2L Both the conventional machining and LAM were performed at different cutting speeds and laser powers at the fixed depth of cut of 1mm and feed rate of 0.214mm per revolution. Chips obtained after machining were mounted with epoxy on their side so that the chip cross-section after polishing was straight across its length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000604_iembs.2008.4650354-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000604_iembs.2008.4650354-Figure8-1.png", + "caption": "Fig. 8 Spastic Walker 2 (with knee and hip flexor support)", + "texts": [ + "9 s, while muscular reflex alone could lead the walker back to the normal walking attractor within 2.9 s. This result suggests that CPG-phase-modulation is effective for modulating the relationship between internal CPG and the body mechanisms enabling a fast recovery to a stable walking cycle. Two walkers, that is, the walker with knee flexor support (Spastic Walker 1) and the walker with knee and hip flexor support (Spastic Walker 2) were applied. As the result of the experiments, Spastic Walker 2 acquired a stable gait pattern. Fig.7 and Fig.8 show the gait of each walker, and Fig.10 shows the hip joint angle phase portraits of each walker. As shown in Fig.8, Spastic Walker 2 lifts its swing leg higher than Normal Walker, and avoids a stumbling. This result is very similar to the result of paralyzed people\u2019s gait experiment [2]. Moreover, it is observed the walker has a peculiar stable limit cycle, which is different from that of Normal Walker (Fig.9 (b)). Five seconds of the Energy Stability Margin of each walker are plotted in Fig.10. It is clear that both knee and hip flexor supports are essential for balance maintenance. Thus it suggests that spastic hemiplegic gait can consist of a part of the function decline, that is pes equinus and the suitable compensated walk for the function decline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003043_icems.2018.8549136-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003043_icems.2018.8549136-Figure7-1.png", + "caption": "Fig. 7. Basic model(2D & 3D FEM model).", + "texts": [ + " The method of increasing the number of parallel circuits at high speed in the method of switching the number of parallel circuits of the motor winding in the high speed operation section does not differ greatly from the actual theoretical method of switching the number of serial turns per phase considerably, but equivalent serial turn Unlike the switching method of numbers, unnecessary parts are not used, and the number of turns of all windings is used. In addition, the efficiency is increased due to the resistance is reduced in the high speed section and the motor loss is reduced. Thus, this approach can be extended the average high efficiency range over the entire operating range. However, in such a parallel branches conversion method, as can be seen from the circuit of fig.4, a thyristor for current control is often used. Therefore, circuit size, weight and price have high problems. Fig.7 shows the 2D and 3D FEM models of the basic model respectively. In the case of a basic model, a high efficiency range is displayed in the high speed operation range. The winding switching method applied in this paper aims to expand the range of efficiency at high speed. Therefore, in the case of the basic model, efficiency of the above conversion method cannot be greatly maximized. Therefore, complementary design of the motor is needed to maximize the effect of the method studied in this paper. Fig.7 shows the 2D and 3D FEM models of the basic model respectively. In the case of a basic model, a high efficiency range is displayed in the high speed operation range. The winding switching method applied in this paper aims to expand the range of efficiency at high speed. Therefore, in the case of the basic model, efficiency of the above conversion method can\u2019t be greatly maximized. Therefore, complementary design of the motor is needed to maximize the effect of the method studied in this paper. TABLE I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000779_icit.2009.4939600-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000779_icit.2009.4939600-Figure7-1.png", + "caption": "Fig. 7. Eddy current distribution in stator core.", + "texts": [ + " Je is calculated using (10) { } [ ] [ ] [ ] { }e Tn i Ae AN nt AJ \u2211 = \u2212= \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 \u2202 \u2202\u2212= 1 1\u03c3\u03c3 (10) where Je is current density component due to magnetic vector potential A, [ ]\u03c3 is conductivity matrix, n is the number of integration points. Js is calculated using (11) { } [ ] [ ] { } { }e Tn i s VN n VJ \u2211 = \u2207=\u2207\u2212= 1 1\u03c3\u03c3 (11) where Js is current density component due to V, \u2207 is divergence operator, Ve is electric scalar potential, N is element function for V evaluated at the integration points. The velocity current density vector Jv is calculated using (12) { } { } { }BvJv \u00d7= (12) where v is applied velocity vector. The distribution of eddy current in stator core is shown in Fig. 7. The flux density loci at typical elements of the laminated steel motor yoke are shown in Fig. 8. Core loss is crucial for flameproof motor because the core loss is the dominant component of power loss. An accurate model including end windings is applied for predicting the core loss in laminated steel flameproof motor. The total core loss is computed based on time-stepping finite element analysis by separating hysteresis, eddy current and anomalous loss in each element under the alternating and rotational magnetic fields" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003069_asemd.2018.8558991-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003069_asemd.2018.8558991-Figure1-1.png", + "caption": "Figure 1. Machine topology.", + "texts": [ + " The function of the DC filed winding can be switched from magnetization manipulation to hybrid excitation, and vice versa, according to the operation. 3D FEA is used to evaluate the impact of the parameters, such as the proportion between two types of PMs and field current density, on the machine performance. A prototype machine is built and tested to validate the predicted results. II. MACHINE TOPOLOGY AND WORKING PRINCIPLE The topology of the proposed dual-stator SPM-AFFSM machine is shown in Fig. 1. Each stator adopts a 12 stator slots(S)/ 10 rotor poles (P) configuration. Two types of PMs of same magnetization direction, i.e. low coercive force (LCF) PM and high coercive force (HCF) PM, are connected in series and embedded into the stators. Separated PMs in each stator are in opposite magnetization directions. Symmetrical PMs in two stators are also in opposite magnetization directions. An additional DC field winding is introduced to the inner side of the stators. The working principle of the proposed SPM-AFFSM is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003238_ciced.2018.8592327-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003238_ciced.2018.8592327-Figure3-1.png", + "caption": "Fig. 3 3d simulation model and subdivision diagram of High frequency carrier line element", + "texts": [ + " The simulation model is shown in figure 2(a), and the segmented grid after meshing is shown in figure 2(b). In the simulation model, the material of the cabinet is steel, and there is a 2mm gap between the cabinet body and the cabinet door for electromagnetic wave to spread out. High frequency current-carrying coil is used to simulate the partial discharge generated after the failure in the switch cabinet. High frequency current-carrying line element is placed in the simulation model of the switch cabinet. The local magnification is shown in figure 3(a), and the segmented grid is shown in figure 3(b). CICED2018 Paper No. 201804190000003 Page3/6 209 Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 16:43:24 UTC from IEEE Xplore. Restrictions apply. Detection impedance is 50KZ = \u2126 , Partial discharge pulse current amplitude 0 100I mA= , when the pulse width is 10ns, it is approximately equivalent to the discharge power of 1000pc at the fault point. As put increasing distance between power supply and switch cabinet metal shell, induction voltage pulse amplitude decreases, and put the distance between power supply and switch cabinet metal shell increased from 1 cm to 20 cm, the change of the voltage pulse as shown in figure 4, and voltage pulse amplitude change with the distance is shown in fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000947_sice.2008.4655222-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000947_sice.2008.4655222-Figure8-1.png", + "caption": "Fig. 8 The ball position in the image has a deviation from the position on the table.", + "texts": [ + " 6 shows the histogram of hues of all the patches. The hue of green cloth is automatically extracted as a range of hue having high frequency in the histogram. By removing the background (the pixels having the hue in the range) from the image, ball areas are extracted as shown in Fig. 7(b). The position of each ball area is, then, calculated by a pattern matching using an ideal ball pattern as a template. The calculated ball position, however, has a deviation from the actual position on the table as shown in Fig. 8. This deviation (x\u2019) can be easily corrected based on the ball distance L from the center of the table (the camera position) because both the camera height (Z) and the ball diameter (r) are known. The corrected ball position is shown as a cross in Fig. 7(c). An experiment was carried out to evaluate the accuracy of ball position detection. The average error of the ball positions was about 2.3mm on the table (about 5% of the ball diameter). This error, we think, is an acceptable level for the beginners" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000012_6.2008-4509-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000012_6.2008-4509-Figure1-1.png", + "caption": "Figure 1 Schematic of a Stationary Brush Seal", + "texts": [ + " Cincinnati, Ohio, 45246 In this paper, we will present the results of the initial rotational tests for the rotating intershaft brush seal. Data will be presented comparing the rotating brush seal, both single and tandem arrangements to multi-tooth labyrinth seals and also fracture ring carbon seals. I. Introduction With their increasing use in gas turbine engines, brush seals have drawn significant attention in recent years. In particular, brush seals have consistently demonstrated superior performance compared to labyrinth seals. For a typical, air-to-air leakage path, Figure 1 the brush seal consists of fine diameter metallic or ceramic fibers packed between retaining and backing plates. As shown in Figure 1, the backing plate is situated downstream of the bristles to provide mechanical support for the differential pressure loads. The radial seal is attached to a static outer member with the bristles touching inner rotating shaft with an angle in the direction of the inner shaft rotation. This cant angle is a critical parameter as it helps bristles to bend rather than buckle in case of the inner shaft radial excursions. This inherent ability of the bristles to bend and not buckle make brush seals primary candidates for sealing between stationary and rotating parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003394_6.2019-1747-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003394_6.2019-1747-Figure3-1.png", + "caption": "Figure 3: Definitions for a tape spring hinge in deployed (left) and non-symmetric (right) configurations.", + "texts": [ + " The actual force and torque response of the hinge will depend on the loading of either side of the hinge, and small displacements from the nominal configuration may introduce significant force and torque responses. Therefore, the established moment-curvature approach is not sufficient for the modeling fidelity desired here, and a study of force and torque responses due to non-symmetric behavior is conducted. The phenomenon of undesirable non-symmetric configurations in the tape spring hinge fold is not well studied. Here, nonsymmetric behavior refers to any change in position and orientation that does not follow the nominal fold rotation, as is illustrated in Figure 3. To guarantee symmetric behavior, additional components must be included in a hinge assembly to constrain the hinge, which can add mass and complexity where lightweight simplicity is desired. Such solutions are not addressed here. Inclusion of multiple independent state variables in this study makes it difficult to approach the problem with classical theory, therefore, to study this phenomenon, numerical and experimental techniques are employed. High strain composites are a novel class of flexible material with great potential for spacecraft deployable structures", + " A quick study is conducted to observe the maximum principle strains occurring in the tape spring for various lengths when the hinge undergoes a nominal fold to 90 deg using an FEA simulation. The results are shown in Figure 2 for both materials undergoing an equal sense fold, and a length of 150 mm is selected for this study. This is done for a tape spring with 20 mm long clamps attached at each end point, resulting in a shorter effective composite hinge section. The tape spring hinge is represented in the rigid body dynamics simulations as an internal forcing function in terms of the position and orientation of the hinge connection points. This concept is illustrated in Figure 3, where the fixed end points of the hinge are each assigned a reference frame,A0 andA1, the reaction forces from the hinge are denoted N0 andN1, and the reaction moments are denoted asM0 andM1. These mechanics are modeled as functions of the relative position, \u03b4, and orientation of frame A0 with respect to A1. The hinge model is developed to be compatible with a preexisting multi-body dynamics framework based on the Articulated Body Forward Dynamics approach.13 This approach de-constructs a system of linked rigid bodies by defining the interactions across the hinge connecting an outbound body to an inbound body through relative coordinates, and selecting these as the generalized coordinates of the dynamics model", + " The relative orientation \u03b8(A0,A1) contains 3\u22122\u22121 Euler Angles for ease of interpretation and because the second axis, where the 90 degree Euler angle singularity resides, can be oriented with an axis which does not accommodate significant relative deflection. The A1 frame is oriented identically to theA0 frame when the hinge is deployed in the zero energy state. The displacement of the relative hinge frame coordinates, \u03b4, is selected over the relative position, r, to better correlate the physical behavior with the numerical fit. The relation of these vectors is displayed in Figure 3, defined as \u03b4 = r \u2212 ri (2) Then the generalized forces and torques acting at frame A0 are written as a function of the relative coordinates across the hinge frames, in spatial notation, as f0(q) = [ N0 M0 ] = N01 N02 N03 M01 M02 M03 (3) The common assumption for hinge force and torque models is that the force and torque are acting in equal but opposite direction on each of the connected rigid bodies at the connection frames. While a quick free body analysis of Figure 3 verifies this to be true for the force, the moment balance introduces something new. The summation of moments at either frame will require the torque due to the reaction force and the relative position of the frames be included. Therefore, the spatial force at frame A1 can be written in terms of only the force and torque at frame A0 as f1(q) = [ N1 M1 ] = [ \u2212N0 \u2212M0 \u2212 r \u00d7N0 ] (4) Equation 4 indicates that the force and torque applied to the rigid bodies can be determined for both sides of the hinge using a model of only one set of forces and torques", + " Two ATI six-axis force/torque transducers are used at the reference frames on the hinge to directly measure the full force/torque profile. The transducers are calibrated for torque measurements of 500 N-mm with 1/16th N-mm resolution and forces of 50 N in plane and 70 N out of plane with 1/80th N resolution. These sensors are aligned with the hinge such that the measurement frame of the sensor is coincident and orthogonally aligned to the hinge reference frames A0 and A1. The data from these hinges are then transformed into the frame alignments defined in Figure 3. An NI Labview program is used to interface with the transducers through an NI USB-6218 data acquisition card. The hinge configuration is controlled using multiple stepper motors and a SparkFun RedBoard, also interfaced through the Labview program with identical timing. The hinge configuration is not observed through external means, but is derived through the stepper motor count. The stepper motors are controlled using microstepping, with a resolution of 0.225 degrees per step. The left reference point of the hinge is mounted to a cart controlled through a smooth linear rail and the rotation about a01 is controlled by an additional motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002940_icsedps.2018.8536040-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002940_icsedps.2018.8536040-Figure1-1.png", + "caption": "Fig. 1 2-ph Induction Motor 369", + "texts": [ + " It also does not require current transducer which makes this method cost effective. Also this method performs well during the steady as well as transient condition. The ASFC method can be used where high speed application is required. II. MATHEMATICAL MODEL 2-ph. induction motor: A 2-ph phase induction motor consist of two windings i.e. main winding and auxillary winding. The motor used in our scheme is symmetrical induction motor. Therefore number of turns on d-axis is exactly equal to the turns on qaxis. Both the winding are place 90 degrees apart. Fig. 1 shows 2-ph. induction motor. r\u03b1 is the rotor component on \u03b1 axis. s\u03b1 is stator winding on \u03b1 axis. Similarly s\u03b2 , r\u03b2 are stator and rotor component respectively on the \u03b2 axis. For modeling of induction motor in MATLAB, equation are required which are found from the equivalent circuit. Fig.2 and Fig. 3 shows equivalent circuit in stationary reference frame. By applying KVL in the Fig. 2 and Fig. 3. Stator equation in stationary reference frame are given below \u03b1\u03b1\u03b1\u03b1 \u03c8 ssss dt diRv += (1) \u03b2\u03b2\u03b2\u03b2 \u03c8 ssss dt diRv += (2) Rotor equations in stationary reference frame are given below \u03b2\u03b1\u03b1\u03b1\u03b1 \u03c8\u03c9\u03c8 rrrrrr a dt diRv ++== 0 (3) \u03b1\u03b2\u03b2\u03b2\u03b2 \u03c8\u03c9\u03c8 rrrrrr adt diRv 10 \u2212+== (4) The equation of stator flux linkage component and rotor flux linkages component can be obtained as: \u03b1\u03b1\u03b1\u03b1\u03b1\u03c8 rmsss iLiL += (5) \u03b2\u03b2\u03b2\u03b2\u03b2\u03c8 rmsss iLiL += (6) \u03b5\u03b1\u03b1\u03b1\u03b1\u03b1\u03c8 rrrsmr iLLiL ++= (7) \u03b2\u03b2\u03b2\u03b2\u03b2\u03c8 rrsmr iLiL ++= (8) Using equation (9)-(12), the rotor and stator currents equations can be obtained as: \u03b1\u03b1\u03b1 \u03b1\u03b1\u03b1\u03b1 \u03b1 \u03c8\u03c8 mrs rmsr s LLL LLi 2\u2212 \u2212= (9) \u03b2\u03b2\u03b2 \u03b2\u03b2\u03b2\u03b2 \u03b2 \u03c8\u03c8 mrs rmsr s LLL LL i 2\u2212 \u2212 = (10) \u03b1\u03b1\u03b1 \u03b1\u03b1\u03b1\u03b1 \u03b1 \u03c8\u03c8 mrs smrs r LLL LLi 2\u2212 \u2212= (11) \u03b2\u03b2\u03b2 \u03b2\u03b2\u03b2\u03b2 \u03b2 \u03c8\u03c8 mrs smrs r LLL LL i 2\u2212 \u2212 = (12) The electromagnetic torque equation can be given by the equation as: ( )\u03b2\u03b1\u03b1\u03b1\u03b2\u03b2 rsmrsmpe iiLiiLPT \u2212= (13) And the mechanical dynamic is modeled by the equation Ler TT dt dJ \u2212=\u03c9 (14) where \u03b1sv , \u03b2sv , \u03b1rv , \u03b2sv represent stator and rotor voltages, \u03b1si , \u03b2si , \u03b1ri , \u03b2si represent stator and rotor currents, ,s\u03b1\u03c8 \u03b2\u03c8 s , \u03b1\u03c8 r , \u03b2\u03c8 s represents stator and rotor flux linkages \u03b1sR , \u03b2sR , \u03b1rR , \u03b2sR represents stator and rotor resistances, \u03b1sL , \u03b2sL , \u03b1rL , \u03b2sL represents stator and rotor inductances, \u03b2\u03b1 mm L,L represents magnetizing inductances, r\u03c9 represents electrical rotor angular velocity, eT is the electromagnetic torque, LT represents load torque, J is the rotor moment of inertia, dt d is the differential operator and a is the main per auxiliary winding turns ratio [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003843_012039-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003843_012039-Figure4-1.png", + "caption": "Figure 4. Prototype DVA Machine (Dynamic Vibration Absorber)", + "texts": [], + "surrounding_texts": [ + "IOP Conf. Series: Materials Science and Engineering 494 (2019) 012039 IOP Publishing doi:10.1088/1757-899X/494/1/012039" + ] + }, + { + "image_filename": "designv11_92_0002332_speedam.2018.8445270-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002332_speedam.2018.8445270-Figure5-1.png", + "caption": "Fig. 5. Vector of magnetic flux density", + "texts": [ + " FFT analysis result of the back-EMF in the magnetized and demagnetized conditions is presented in Fig. 3. When the motor condition changes from the magnetized condition to the demagnetized condition, the back-EMF can be reduced from 100 % to 60.6 % as shown in this figure. Fig. 4 shows the magnetic flux density in the air gap. In the magnetized condition, the magnetic flux density is about 0.98 T when the magnetic flux linkage is maximum. On the other hand, in the demagnetized condition, the magnetic flux density is about 0.58 T. Fig. 5 (a) and Fig. 5 (b) depict the vector of the magnetic flux density in the magnetized and demagnetized conditions. In the stator, we can observe that the magnetic flux is increased in the magnetized condition and is decreased in the demagnetized condition from these figures. The torque versus current phase angle characteristics in both conditions when the U-phase current is 0 are shown in Fig. 6 (a) and Fig. 6 (b), respectively. As is obvious from Fig. 6 (a), when the current phase angle is 40 \u00b0 in the magnetized condition, the torque shows the maximum value of 76" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002046_978-3-319-79005-3_17-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002046_978-3-319-79005-3_17-Figure5-1.png", + "caption": "Fig. 5 Three-dimensional model", + "texts": [ + " The specimen model was generated with standardized measurements established by ASTM D638 for tensile tests on flat test specimens (Fig. 4). The standard establishes the following dimensions for the specimen: W Width of the narrow zone 6 mm L Length of the narrow zone 33 mm WO Width of ends 19 mm LO Total length 115 mm G Calibrated length 25 mm D Distance between jaws 65 mm R Radio 14 mm T Thickness 4 mm Defined the dimensions of the specimen the model was generated. In this case, the Solidworks 2015 design program was used for this purpose (Fig. 5). The part was modeled in SolidWorks, and exported with extension (IGS) to CAM software, which is connected to the daVinci 1.0 printer with FDM technology. Before Fig. 6 ABS Printed specimens Fig. 7 Filament directions 45\u00b0 and 135\u00b0 starting the printing of the specimens, it was necessary to heat the injector at 230 \u00b0C and heat the printing base to 110 \u00b0C for the ABS. Since the printer works with interchangeable layers of 90\u00b0 of separation, the 3D prints (Fig. 6) were made with filament orientation of 45\u00b0 and 135\u00b0 (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002607_012043-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002607_012043-Figure8-1.png", + "caption": "Figure 8. Finite element model of a Qloop bonding wire.", + "texts": [ + " Second, the gap between parallel adjacent wires and corresponding mechanical shock load, which are also listed in the table 3 and will be used to calculate the damping coefficient. 3.1. Extraction of characteristic parameters of bonding wire This paper use loop span, loop diameter and loop height to describe a Q-loop bonding wire. Q-loop is a simple loop profile which is defined in the wire bonder and its structure is shown in figure 7. Different finite element models are established by changing the three structure parameters and the structure is shown in figure 8. This paper selects multiple parameters nodes in order to get the relationship between structure parameters and vibration parameters. The range of loop span, wire diameter and loop height is shown in table 4 and the sample points are also shown in it. 4 because of its widespread use. 3.2. Vibration analysis of bonding wire 3.2.1. Initial velocity calculation. The initial velocity of bonding wire should be calculated by the experimental data before transient dynamic analysis. The mapping relationship between initial amplitude and mechanical shock load can be built from the data, which is shown in figure 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001071_0263-0923.28.4.269-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001071_0263-0923.28.4.269-Figure1-1.png", + "caption": "Figure 1. Motor coordinate system", + "texts": [ + " In the second section, a new sensorless motor parameter identification method is proposed. In the third section, a new sensorless control that does not require the assumption that the motor speed is constant is proposed. In the fourth section, the effectiveness of the methods are verified using numerical simulations with the implementation of a Pulse Width Modulation (PWM) inverter. A new motor parameter identification method is proposed here, and it is presented that the method is available without using information as to the rotor angle. The motor coordinate system is shown in Fig. 1. \u03b8 is the rotor angle. The subscripts U, V , W mean quantities on a fixed 3 phase axis, \u03b1, \u03b2 mean quantities on an orthogonal fixed 2 phase axis, and d, q mean quantities on an orthogonal 2 phase axis synchronized with the rotor angle \u03b8 [2]. \u03b3, \u03b4 mean quantities on the orthogonal 2 phase axis synchronized with estimated rotor angle \u03b8\u02c6 derived by sensorless drive. \u03b8e is yielded by (1) dq-axis corresponds with \u03b3\u03b4-axis when \u03b8e =0. Parameters Identification for Interior Permanent Magnet Synchronous Motor Driven by Sensorless Control JOURNAL OF LOW FREQUENCY NOISE, VIBRATION AND ACTIVE CONTROL270 The IPMSM model on the \u03b1\u03b2 fixed coordinate is yielded by the following equations [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002193_978-3-319-99262-4_31-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002193_978-3-319-99262-4_31-Figure2-1.png", + "caption": "Fig. 2. Schematic view of the square dimples [11]", + "texts": [ + " A Cartesian coordinate system is also shown in Fig. 1. The x-axis is along the vertical direction, the z-axis is in the direction of the journal length, and the Y axis is perpendicular to the x- and z-axes. The circumferential coordinate h is measured from the positive X axis in the rotating direction. An axial oil groove is set on the top of the bearing and lubricating oil is supplied from it. The texture region is formed from hts to hts+ hs on the bearing surface in the circumferential direction, and to the whole in the axial direction. Figure 2 shows a schematic of the dimples on the bearing surface. We chose the square dimples according to our previous study [11, 12]. As shown in Fig. 1, the square dimples are formed at an equal interval in the texture region. Three parameters characterize the square dimples; the depth hd, the length and the pitch. Considering the direction, the length are represented ldh, ldz in the circumferential and axial direction respectively, and the pitch lph, lpz in the circumferential and axial direction respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003610_joe.2018.9382-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003610_joe.2018.9382-Figure1-1.png", + "caption": "Fig. 1 Flat voice coil actuator", + "texts": [ + " The flat VCA is designed according to the theory of Lorentz force principle. It consists of a magnet and a coil, with the magnet fixed to the PM and the coil fixed to the SM of the non-contact ultraquiet satellite. The magnet frame is used to install permanent magnet to produce permanent magnetic field and the coil frame is used to circle coils. When the coils are connected with electric current, the Lorentz force would be generated, and the movable magnet frame will move vertically and horizontally. The Lorenz force principle of the VCA IS shown below. According to Fig. 1, the coordinate system of the magnet is defined as O1X1Y1Z1, and the coordinate system of coil is defined as O2X2Y2Z2. The direction of the magnetic field strength B keeps along with O1Y1, and the current direction keeps along with the axis O2X2. It could be found that the area of the coil covered by the permanent magnet in the range of motion remains unchanged, ensuring that the output force of the VCA is mainly dependent on the current, and having little to do with the relative position of the mover, that is, the high precision of the output force is ensured", + " The eddy current position sensor and the flat VCA are designed as an integrated structure, with the probe fixed on the magnet through the bracket. It has the advantages of high-precision relative position measurement and high-precision force output. The structure is also compact and easy to install. The integrated structure of the eddy sensor and the voice-coil actuator is shown below (see Fig. 3). A mathematical model of output force of the flat VCA is established. In the coordinate system, shown in Fig. 1, Euler angle is adopted to describe the relative attitude parameters of O1X1Y1Z1 and O2X2Y2Z2. Assuming that O1X1Y1Z1 rotates to O2X2Y2Z2 in 3-2-1 order, the attitude matrix between O1X1Y1Z1 and O2X2Y2Z2 could be denoted as (see equation below) where \u03c8 , \u03b8, \u03c6 is attitude angle. The magnetic field intensity direction of the VCA is fixed to the coordinate O1X1Y1Z1, and thus, the coordinate could be described as B1 = Buy; the conductor direction is fixed to the coordinate O2X2Y2Z2, and its coordinates are L2 = lux; the current is I, and the direction is the same as the conductor direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003115_ecce.2018.8557727-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003115_ecce.2018.8557727-Figure1-1.png", + "caption": "Fig. 1. Cross section of a triple redundant, 9-phase PMASynRM. (a) Named slots and short-circuit turn. (b) Winding Layout.", + "texts": [ + " Firstly, a 3D transient thermal model and a LP thermal model will be established. Then the temperatures predicted by the two models under the same fault conditions will be compared to validate the LP model. Moreover, the transient and steady-state temperatures predicted by the 3D thermal model considering more realistic issues will be comprehensively compared with the experimental results. The merits of the two thermal models are also assessed. II. 9-PHASE (3X3-PHASE) PMASYNRM A triple redundant, 9-phase (3x3-phase), 36-slot, 6-pole PMASynRM as shown in Fig. 1 is reported in [11]. This machine has comparable performance with conventional PM machines in terms of efficiency and torque density due to high saliency and inherent large reluctance torque. It also has low permanent magnet usage which results in the low back-EMF and low short circuit current, both of which improve the fault tolerance. Fig. 1 (b) shows the layout of windings. Compared with the conventional overlapped distributed windings, this machine employs three separated 3-phase windings which do not overlap with each other. They are denoted as ABC, DEF and GHI. This winding layout improves the physical and thermal isolations between the different 3-phase sets. Apart from that, three standard 3-phase inverters are used to drive each 3-phase set to provide the electrical isolation. Thus, fault propagation between different 3-phase winding sets is minimized", + " Without loss of generality, it is assumed that the worst case one turn SC occurs in phase B and terminal SC will be applied to 3-phase set ABC when the fault is detected. The mutual coupling between the two healthy 3-phase sets and one faulty 3-phase set will result in an magneto-motive force (MMF) offset component in the region occupied by 3-phase set ABC [14]. Therefore, the location of the SC turn will affect the flux linkage and consequently the circulating current and resultant copper loss. It has been investigated in [15] that when the SC turn located at slots B2 and B4 which are marked by the two black quadrangles shown in Fig. 1 (a), the SC current and copper loss are the highest. The subsequent analysis is focused on this worst case. III. 3D THERMAL MODEL The 3D thermal model is shown in Fig. 2 where different components are indicated. The 3D model encompasses 12 slots and one half of the machine axial length. To reduce modelling complexity, the end winding is simplified in the 3D thermal model as straight winding segment with the same equivalent length as those in the prototype machine. The 3D FE model together with the schematic heat transfer network with the ambient and cooling circuit shown in Fig", + " COMPARISON OF 3D THERMAL MODEL AND TEST The 3D thermal model would be more accurate since it can predict transient temperatures and cope with a number of practical issues compared with the LP thermal model. The first practical issue is that the inlet, outlet temperatures and flowrate of cooling oil may change during operations. This condition can be modelled in 3D thermal model by setting the ambient temperature boundary and the thermal convection resistance as time-dependent functions. In addition, the end winding layout of a real machine is non-uniform as shown in Fig. 1 (b) with the middle part having more conductors than the two sides in each 3-phase set. To represent this effect, the copper loss density in the end winding conductors associated with different slots is set proportional to the real winding layout. Moreover, the copper loss is dependent on temperature and EM-thermal coupled simulation considering temperature effects is the most accurate to address this problem. However, in this study, currents are considered to be independent of temperature. Therefore, the EM-thermal coupled simulation is not necessary, but copper loss variation with the winding temperature is represented in the 3D thermal model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002839_s1028335818100105-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002839_s1028335818100105-Figure2-1.png", + "caption": "Fig. 2. Stability region of the inverted pendulum. N(0.3054, 0.9368), , 0.8430), and ; points and are the same as in Fig. 1.", + "texts": [ + " At point , mapping (6) is reduced to form (11), in which where . Since , the immobile point of mapping (6) is stable [10, 11]. Thus, the problem of the hanging-pendulum stability is solved for all values of parameters a and b that lie in region (8) or at its boundaries. (Recall that the calculations were carried out for ). Omitting the details, we report only the results obtained. The calculations were performed for . In the first approximation, the stability region under study is located (see Fig. 2) between curves \u03c7 = \u2212 +2 2 3 3 3 3| |c d e 3 3, ,c d 3e 4F 6F 3P 6P \u03c73 \u03c73 3P \u03c73 6P 3P 6P 4 \u03b3 ( , )D E \u03bb =5 1 \u03bb =5 2 \u03b3 \u03b3 \u03c0\u03be = \u03be + \u03be + \u03b7 \u03b7 + \u03c0\u03b7 = \u03b7 \u2212 \u03be + \u03b7 \u03be + 2 2 2 3 7 2 2 2 3 7 3 ( ) and* 2 3 ( ) ,* 2 \u0441 O \u0441 O 3\u0441 \u22603 0\u0441 3\u0441 \u03b3 *P *P \u03be = \u03be + \u03c0 \u03be + \u03b7 \u03b7 + \u03b7 = \u03b7 \u2212 \u03c0 \u03be + \u03b7 \u03be + 2 2 3 4 9 2 2 3 4 9 ( ) ,* ( ) ,* \u0441 O \u0441 O =4 17.4233\u0441 \u22604 0\u0441 \u2264 \u22640 10a \u2264 \u22640 1a \u03b3r DOKLADY PHYSICS Vol. 63 No. 10 2018 and . We have and at the boundary curves and , respectively. The curve of the fourthorder resonance is shown by the dash-dotted line in the stability region. The portions of the boundary curves at which the inverted pendulum is stable and unstable are shown in Fig. 2 by the dashed and solid curves, respectively. The instability occurs at point M and stability occurs at point N. In the first approximation, stability occurs for the entire stability region, except for the half-interval of the curve , in which the inverted pendulum is unstable. This study was performed within state contract no. AAAA-A17-117021310382-5 and supported in part by the Russian Foundation for Basic Research, project no. 17-01-00123. 1. A. A. Andronov, A. A. Vitt, and S. E. Khaikin, Theory of Oscillations (Fizmatgiz, Moscow, 1959) [in Russian]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002316_1.5051140-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002316_1.5051140-Figure1-1.png", + "caption": "Fig. 1. The case of an object acted on by gravity and a velocity dependent drag force. Velocity of the object is also shown.", + "texts": [ + " Furthermore, it would be impossible to show that the fall times are exactly the same; one could only set an upper limit on the time difference. It occurred to us that there is another form for the drag force that is straightforward to examine both theoretically and in the lab. That is the case of a constant magnitude drag force (one dependent on the zeroth power of the objects\u2019 speed). For example, this occurs when the drag force is due to sliding friction. That situation is the subject of this paper. Objects falling under the influence of gravity and drag force: Three cases Consider the situation illustrated in Fig. 1. An object with instantaneous velocity v is affected by gravity Fg and experiences a drag force FD with a direction opposite that of the velocity. The magnitude of the net vertical force Fnet(y) on the object is given by: Fnet(y) = FD cos \u03b8 \u2013 Fg. (1) It is also evident from Fig. 1 that the magnitude and components of the velocity are given by vx = v sin \u03b8, vy = v cos \u03b8, and thus (2) We consider three possible drag forces with a magnitude The Physics Teacher \u25c6 Vol. 56, September 2018 341 exerts on an object with mass m is FN = mg cos \u03c6 and the object is forced down the incline by a component of gravity having a magnitude mg sin \u03c6. The normal force acting on the object is FN = mg cos \u03c6. (9) Another view of an object that follows a curved path along an inclined plane is illustrated in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002497_978-3-319-99620-2_12-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002497_978-3-319-99620-2_12-Figure3-1.png", + "caption": "Fig. 3. Model of 1/4 of the vessel and finite element mesh used in FEA.", + "texts": [ + " Crack occurrence in this zone may lead to a complete failure of the vessel, as shown in [6, 27]. Figure 2b displays the strain distribution along the part of the weld joint of the vessel head. Value of the strain in that area varies from 0.10 to 0.25%, with dominant value of approximately 0.15%. Based on the dimensions of the pressure vessel used in experiments, geometry of the vessel needed for numerical simulation was made in CATIA v5 software. Due to the symmetry of the vessel, only one quarter was designed and exported to Ansys Workbench software (Fig. 3). Applied boundary conditions and loads matched those used in experiments, while several different mesh densities had been used until the most appropriate was obtained. Final mesh consisted of 145,701 nodes and 72,732 solid elements. After applying pressure load, static nonlinear FEAs were carried out using material data obtained previously in the experiment with specimen [28]. The main purpose of these numerical simulations was verification of the developed model and \u2013 as it can be seen in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000569_isma.2008.4648827-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000569_isma.2008.4648827-Figure1-1.png", + "caption": "Figure 1. Phase plane portrait showing SMC action", + "texts": [ + " The stability of a SMC designed using the proposed methodology, and comparisons with proportional controllers are provided. The design methodology explained in [14] is suitable for Single-Input-Single-output (SISO) control systems. Hence, the design methodology in this paper is extended to deal with SIMO control systems. An inverted pendulum system is selected to test the new design methodology. The system is a special case of Multi-Input-Multi-output (MIMO) control systems and is acknowledged to be difficult one to control. Figure 1 shows a phase portrait of a system moving form point A to the equilibrium point at 0 along the shown sliding surface. A controller is required to move the system to the desired point at 0. Applying the principle of work and energy results in the following relationship between the controller gain and the slope of the sliding surface shown in Figure 1 where the controller is a relay type controller [14]: \u03b1\u03b1\u03b2\u03bb \u0393++= )( 2 1 2k (1) where, \u03b2 is the inertial energy storage element constant \u03b1 is the potential energy storage element constant \u0393 is the linear dissipative element constant k is the controller gain \u03bb is the slope of the sliding surface shown in Figure 1 To find the parameters of Equation 1, consider a mechanical system where m, b, and sk are the mass, damping coefficient, and spring constant respectively of the system. With reference to [14], it was shown that: ,, skm \u2261\u2261 \u03b1\u03b2 and b\u2261\u0393 Equation 1 can be used for SISO control systems. An extended design methodology can be achieved to include SIMO control systems. If a four-state-variables system with a single input is considered, two sliding surfaces can be designed as follows: 2111 xxs += \u03bb (2) 4322 xxs += \u03bb (3) where 1x , 2x , 3x and 4x are the states of the system and 12 xx = , 34 xx = " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002058_i2mtc.2018.8409547-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002058_i2mtc.2018.8409547-Figure4-1.png", + "caption": "Fig. 4 The environment of the experiment: (a) SQI mechanical failure simulator and (b) the position of the sensors", + "texts": [ + " For the lack of space, we just give the experienced values =0.25 , =0.001 with 1 4N , 5M . IV. EXPERIMENTAL AND ENRINEERING DATA VALIDATION In this section, the proposed method is utilized to diagnose the fault of a motor bearing. The experiment was carried out on a Spectra Quest Inc (SQI) mechanical failure simulator, which can simulate faults of motors, rolling bearings, shafting, gears and so on. This device controls the output speed of a motor through a speed controller, and then controls the belt drive and gears. The simulator is shown in Fig.4(a), and the sensors are mounted on the motor as shown in Fig.4(b). In this experiment, we mainly diagnose a localized defect on the outer race of the motor bearing. Signals are measured by an accelerometer and Sony EX data acquisition system. Specifically, the type of the motor bearing as SKF6203, the sampling frequency is 6400Hz, and the output speed of the motor is 1433r/min. According to the size of the bearing and rotating speed, we can calculate the characteristic frequency of outer race 73.2Hzof . One component of measured signals sustaining one second is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001102_08ias.2008.178-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001102_08ias.2008.178-Figure9-1.png", + "caption": "Fig 9: (a) Block diagram of the experimental system, (b) Photograph of the laboratory experimental setup.", + "texts": [ + " For simulation the parameters of PI and NFC are listed as follows: PI: , , , NFC: Learning rate = 0.1, Initial values 1~9 2.5w = \u2212 , 19~ 27 2.5w = + The parameters of the FIM are shown in Appendix. The motor was damaged by introducing a rotor fault as shown in Fig. 1. The proposed self tuned NFC based vector control of FIM drive system has been implemented in real-time using the DSP board DS1104 [17] for a laboratory 250 W faulty IM with broken rotor bars11. The block diagram and the photograph of the experimental system are shown in Figure 9 (a) and (b), respectively. The PC-based controller produces numerical switching commands sent to DSP board and the outputs of the DSP board are sent to the amplifier circuit to drive the VSI inverter. The actual motor currents are measured by the Halleffect sensors and fed back to the DSP board through the A/D channels. The rotor position is sensed by an optical incremental encoder of 1000-line resolution and is fed back to the DSP board through the encoder interface. The test FIM is coupled to a dc machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.4-1.png", + "caption": "Figure 1.4 The ball-on-disc machine.", + "texts": [ + " Correction formulas are available [5] for these thermal and starvation effects. 1.4 EXPERIMENTAL ELASTOHYDRODYNAMICS The experimental investigation of EHL benefits from a geometry that, compared with say a journal bearing, is relatively easy to access so that the two important parameters of film thickness and traction may be measured in a bench-top setting. Two embodiments of what are known as EHL rigs are illustrated in Figs. 1.4 and 1.5. The most common today is the ball-on-disc machine shown in Fig. 1.4 where a steel ball is loaded against a disc which may be fabricated from steel or a transparent material such as glass or sapphire. When the disc is transparent, it is a simple matter to measure the film thickness of the liquid between the ball and disc by interferometry. However, there is a large variation in the accuracy of the various interferometry techniques. The twin-disc machine shown in Fig. 1.5 is better suited to traction measurement than film thickness; however, this configuration has been used for film thickness measurement as well by the electrical capacitance method. The shear force transferred across the film is known as traction or friction interchangeably. Traction in the rolling direction, longitudinal traction, may be measured as the force shown as the bold arrow in Fig. 1.4. The component of longitudinal traction due to shear of the liquid film results from a difference in velocities of the surfaces of the ball and disc in the rolling direction, say u1andu2. The rolling velocity is u5 u11 u2\u00f0 \u00de=2 and the longitudinal slide/roll ratio is \u03a3L 5 \u00f0u12 u2\u00de=u. A smaller portion of the longitudinal traction results from the Poiseuille shear in the inlet zone and is present even in pure rolling, \u03a3L 5 0. The twin-disc machine is also convenient for the measurement of the transverse traction that results from a transverse slide/roll ratio, \u03a3T 5 \u00f0v12 v2\u00de=u, where v is the component of surface velocity normal to the rolling direction. The transverse sliding condition can be imposed by skewing the rollers by an angle shown in the upper view of Fig. 1.4 and the associated transverse traction can be measured as the thrust indicated by the bold arrow in the lower view. Traction is generally reported as a traction coefficient, the ratio of traction force to normal force or load, W. Equivalently, the traction coefficient is the ratio of average shear stress to average pressure, \u03c4=p. An additional kinematical condition is spin, the relative rotational velocity of the surfaces. Spin has been eliminated from the contact between the ball and disc in Fig. 1.4 by having the rotational axis of the ball intersect the axis of the disc at the contact plane. Other geometries may involve spin. Spin is generated by a nonzero tilt angle shown at the bottom of Fig. 1.5. Some measurements of central film thickness are shown in Fig. 1.6. The data points have been reported by INSA, Lyon [12 14] using their colorimetric interferometry technique that has shown to be very accurate and repeatable. The contact pressure here was low, pH 0:5 GPa, generated by the contact between a steel ball and glass disc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003507_elk-1702-150-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003507_elk-1702-150-Figure5-1.png", + "caption": "Figure 5. Prototype photograph of the SIDM.", + "texts": [ + " The indicator function of NNC is expressed as follows: J2 = [(r \u2212 y)\u2212\u2206y\u0302] 2 \u2264 \u03b52 (12) The parameters of the NNC are adjusted with the gradient decent algorithm, whose detailed expressions are as follows: wpd(k) = [wpd1(k);wpd2(k)]), (13) dwpd(k) = a2 \u00d7 ((r(k)\u2212 y(k))\u2212\u2206y\u0302(k))\u00d7 [ek; ek \u2212 ek\u22121], (14) wpd(k + 1) = wpd(k) + dwpd(k), (15) where the a2 is learning rate for the NNC. An experiment is performed to validate the availability of the designed controller. The prototype of the designed SIDM is shown in Figure 5. The actuator mainly includes a piezoelectric stack, a frictional rod, a slider, and a preload mechanism. The piezoelectric stack (AE0203D08DF, NEC-Tokin Corporation, Japan) is fixed tightly and the d33 working mode is adopted. The frictional rod is bonded to the piezoelectric element by epoxy resin adhesive. Carbon fiber reinforced plastic (CFRP) is used for the frictional rod due to its superior merits of high stiffness and low density. The slider is preloaded to the frictional rod with the preload mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003231_icsai.2018.8599363-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003231_icsai.2018.8599363-Figure6-1.png", + "caption": "Figure 6. The large stealthy air aircraft carrier model architecture", + "texts": [ + " In Figure 5, the interception net screen is made of the extremely strong transparent silk thread such as nylon wire. The mesh can be customized according to size, and the smallest mesh can be used to catch small birds flying in the sky. III. MODEL ARCHITECTURE OF LARGE STEALTH AIRCRAFT It is necessary to prevent the possible attacks of swarm UAVs to ensure the safety of urban public infrastructure and commercial flights. Therefore, it is necessary for the large stealth aircraft to have the ability of carrying stealth UAVs to maintain the safety of the air control system and the safety of the city. Figure 6 shows the system model of a large stealth fighter with the stealth UAVs. This model is based on the improved model of the Russian Tupolev PAK-DA (Prospective Air Complex for Long Range Aviation) Future Strategic Stealth Bomber model, and also is an improved model of the general model of the stealth UAVs. The principle of the large stealth fighter that response to the swarm of UAVs is to actively release the interception net carried by dual UAVs, two UAVs on its airframe. This is similar to the large passenger aircraft as shown in Figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure5-1.png", + "caption": "Fig. 5 3D model of the rotor-bearing system in a water ring vacuum pump Fig. 6 Results of comparison of the first bending mode for the rotor bearing system. (a) Using the spring element-based approach; (b) using the plate theory-based approach", + "texts": [ + "2 Application example for the proposed method A numerical example of a water ring vacuum pump is Table 1 Calculation of axial and radial deformation parameters Test Modulus of elasticity, E/GPa Axial or radial load, Fa or Fr/N Maximum axial displacement, wmax/mm Radial displacement, wr/mm \u03b1/mm \u03b2/mm Fig. 4 Application procedure for the proposed method [27] studied to validate the effectiveness and efficiency of the proposed model [27]. The purpose is to calculate the first bending mode. The pump is one of the products of a company in Shanghai. The parameters are from a real pump product. As shown in Fig. 5, the pump has two double-row, conical roller bearings, and the outer and inner diameters of the bearings are 210 and 130 mm, respectively. According to Hao et al. [29], the radial stiffness of the drive end bearing is Kr1 = 1.188\u00d7109 N/m, and the radial stiffness of the bearing between the impeller and rotor is Kr2 = 1.365 109 N/m. In this example, the spring element-based method and the plate theory-based model are used to calculate bearing stiffness; the corresponding results are shown in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002171_978-3-319-49574-3_12-Figure12.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002171_978-3-319-49574-3_12-Figure12.5-1.png", + "caption": "Fig. 12.5 Spring mechanism \u2013 ready to open, close, and open", + "texts": [ + " To ensure the voltage is shared across all of the interrupters evenly, grading capacitors can be fitted across each interrupter. A number of different types of mechanism are available. The most common today is the spring type where the opening and closing strokes of the circuit breakers are performed by releasing charged springs. These types of mechanism, if auxiliary power is lost, are able to trip, close, and trip again should the need arise. Should a further close and trip cycle be required, then a hydraulic or pneumatic mechanism that has more stored energy will be required (Fig. 12.5). When the circuit breakers opens, the resultant effect depends upon whether the network is highly interconnected or not. When a fault occurs on the network, an arc from the energized circuit flashes to ground or to another of the energized phases. On detection of the fault, the circuit breakers is commanded to open. If the fault is transient, e.g., a lightning strike, then reclosing of the circuit would result in the circuit being permanently reinstated. Should the fault be permanent, e.g., a fallen tree across the circuit, then this reclosing would be followed by another trip of the circuit breakers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002670_chicc.2018.8483423-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002670_chicc.2018.8483423-Figure1-1.png", + "caption": "Fig. 1: Dynamic Model", + "texts": [ + " In Section 4, simulation results are shown by five missiles salvo attack. Section 5 concludes the full paper and prospects the work in the end. In this section, a leader-follower strategy is adopted to deal with the cooperative guidance for multi-missile. The motion state of leader relative to the target can be considered as the reference input for the tracking control of the follower relative to the target motion state. The planar engagement scene of i-th follower missile, leader missile and target are illustrated in Fig.1. where and are leader missile and follower missile respectively, are the relative distance between target and or , are the line-of-sight angle of or , and are the lead angle of or , and are the angle of velocity of or , and are the normal acceleration of or , and are the velocity of or , and assuming that all the missiles are at the same speed. Assuming the target is static, the state information of leader missile is independent from the followers and target and moves with constant speed. The planar -target relative motion equations are described by: 0 0 0 0 0 0 0 0 0 0 0 0 0 cos sin = + = / r V r q V q a V (1) By adopting proportional navigation guidance (PNG) strategy, the formal acceleration of leader missile can be expressed as follows: 0 0 0 0a = n V q (2) where is the constant coefficient of PNG law" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003742_apsipa.2018.8659673-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003742_apsipa.2018.8659673-Figure7-1.png", + "caption": "Fig. 7. Block diagram of HN (z) with direct-form I structure.", + "texts": [ + " To avoid such overflow at the feedback loop, we apply the following two methods: one is coefficient scaling [13], the other one is direct-form I structure. The coefficient scaling method using Lp norm of the filter can suppress excessive amplification in the feedback loop. However CPZ-ANFs are adaptive filters of which coefficients are time-variant. Accordingly, calculation of Lp norm and division are imposed on DSP and thus the coefficient scaling method is complicated and difficult to implement on DSP. Therefore, we make use of direct-form I structure to implement CPZ-ANFs on fixed-point DSP. Fig. 7 shows the block diagram of HN (z) with direct-form I structure. A difference between direct-form I and direct-form II structure is calculation order of feedforward block and feedback loop. While direct-form II structure first calculates the output of the feedback loop, direct-form I structure first calculates the output of the feedforward block. Overflow at the feedback loop can be avoided by implementing feedforward block first because the feedback loop causes very large gain as mentioned above" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002606_3266037.3271639-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002606_3266037.3271639-Figure3-1.png", + "caption": "Figure 3. An implemented OptRod", + "texts": [ + " The created image is projected to the bottom surface of control unit of OptRod by a light source like a projector or LCD. A phototransistor (PTR) in each control unit senses the brightness of the projected light and Analog to Digital Converter (ADC) converts the brightness into control signal. An obtained control signal is transmitted to the actuator unit through serial bus. A Micro Controller Unit (MCU) in the actuator unit controls an actuator according to the control signal. The implemented prototype of OptRods are shown in Figure 3. The actuator unit is equipped with an actuator (e.g. full-color LED, pin actuator, transducer, and thermoelectric element), a connecting terminal to the control unit, a photo reflector, and a MCU. By using the mounted photo reflector, each OptRod can detect the user\u2019s touch interaction. This feature enables system with OptRods to behave not only as a display that presents information to the user but also as an interactive display. The control unit consists of a connecting terminal to the actuator unit, a power supply terminal, PTR, an infrared (IR) transceiver and an ADC" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002281_01411594.2018.1506127-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002281_01411594.2018.1506127-Figure2-1.png", + "caption": "Figure 2. Caman-like molecules built in the tilted smectic layer: the orientation of c axis in (a) is opposite to that in (b).", + "texts": [ + " Minimization of the free energy (1) with respect to the distribution function f (c) yields the following expression for the equilibrium distribution function: f (c) = 1 I exp \u2212UMF(c) kBT { } where I = \u222bp \u2212p exp \u2212UMF(c) kBT { } dc, (3) where UMF(c) is the mean field acting on a molecule in some tilted smectic layer with dipole moment having orientation \u03c8: UMF(c1) = r \u222bp \u2212p f (c2)Uij(c1,c2) dc2. (4) Substituting Equations (3)\u2013(5) back into Equation (2), one obtains the following expression for the equilibrium polarization-dependent free energy density of the helical smectic liquid crystal: Fp = \u2212rkBT ln I \u2212 1 2 r \u222bp \u2212p f (c)UMF(c) dc. (5) The origin of piezoelectric polarization can be understood easily from comparison of Figure 2(a, b), where the caman-like molecules are built in the tilted smectic layer. Indeed, the orientations of molecules presented in these two figures are different geometrically, and thus, they should correspond to different values of the mean field (5). Due to equal probability of a and \u2212a and also because of the geometrical equivalence of each pair of molecular orientations presented either in Figure 2(a) or (b), the only axis that can have polar ordering is c = [a\u00d7 b], which is perpendicular to both a and b axes. Moreover, the orientation of c axis in Figure 2(a) is opposite to that in Figure 2(b). The simplest intermolecular potential term reflecting the shape of the caman-like molecules can be written in the following form: Upiezo 12 cp(a1 \u00b7 u12)(b1 \u00b7 u12) + cp(a2 \u00b7 u12)(b2 \u00b7 u12), (6) where cp is the piezoelectric constant reflecting the extent of molecular bent related to the camanshape of the molecules, u12 = r12/|r12| is the unit intermolecular vector, a1 and a2 are the long molecular axes, while b1 and b2 are the short molecular axes in the directions of caman nozzles. Let us average potential (6) over the orientation of intermolecular vector u12 within the smectic layer plane having in mind the rule \u3008(u12)a(u12)b\u3009 = (dab \u2212 ka kb)/2, where \u03b1 and \u03b2 both can be one of the two orthogonal frame axes within the smectic layer plane, dab is the Kronecker symbol and k is the smectic layer normal: \u3008Upiezo 12 \u30092 \u2212cp(a1 \u00b7 k)(b1 \u00b7 k) \u2212 cp(a2 \u00b7 k)(b2 \u00b7 k)", + " Amolecular-statistical theory describing the origin of polarization in Sm-C\u2217 and Sm-C\u2217 A is proposed. The Boltzmann distribution for the molecular transverse dipole moments in the mean molecular field is considered. In the framework of this approach, the polar order basically arises due to the bent-shape (caman-shape) of the molecules and can be non-zero even if the molecules are achiral [see Figure 6(a) at mef = 0]. This polar order is related to the energy inequivalence of the camanlike molecular orientations presented in Figure 2(a, b). However, in the absence of transverse dipole moment along the molecular c = [a\u00d7 b] axis, the electric polarization will not arise. In the presence of dipole moment along c axis the molecule becomes chiral and piezoelectric polarization P = rmp arises in each smectic layer. The distribution of charges along the principal molecular axes provides the effective molecular quadrupole, since the molecules are electrically neutral and a and \u2212a orientations are equivalent. The interaction of the molecular transverse dipole with the effective longitudinal quadrupoles in the neighboring smectic layers provides the flexoelectric effect, which is also described in terms of the Boltzmann distribution and results in the helical rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003420_oceans.2018.8604727-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003420_oceans.2018.8604727-Figure4-1.png", + "caption": "Fig. 4. Forces fi \u2200i=[1...4] generated by the four thrusters of the vehicle. CG denotes the position of gravity\u2019s center.", + "texts": [ + " The movement in horizontal plane is referred as surge (along xb axis) and sway (along yb axis), while heave represents the vertical motion (along zb axis). Roll, pitch, and yaw, denoted (\u03c6, \u03b8, \u03c8), are the Euler angles describing the orientation of the vehicle\u2019s body fixed frame with respect to the earth-fixed frame (OI , xI , yI , zI), while (x, y, z) denote the coordinates of the body-fixed frame center in the earth fixed frame. The propulsion system consists in four thrusters, as depicted in Figure 4, which generate the rotational and translational motion. With reference to the rotational movement of this prototype, roll motion is performed through differential speed control of the thrusters 1 and 2. Yaw motion is obteined similarly using thrusters 3 and 4, finally pitch motion is unactuated. On the other hand, the translational movement of the z axis is regulated by decreasing or increasing the combined speed of thrusters 1 and 2. Similarly, the translational movements along the xb and yb axes are obtained by using thrusters 3, 4 and by controlling the yaw angle", + " Based on the design of the vehicle and in order to reduce further analysis, the origin of the body fixed frame is chosen in the gravity\u2019s center, this implies that rw = [0, 0, 0]T ; while the center of buoyancy is rb = [0, 0,\u2212zb]T . For practical purposes, the buoyancy force is greater than the weight, i.e., W \u2212 B = \u2212fb. Notice that fb should be smaller than the force produced by the thrusters. Then from equations (6) and (7), we have: g(\u03b7) = [ fg mg ] = fbsin(\u03b8) \u2212fbcos(\u03b8)sin(\u03c6) \u2212fbcos(\u03b8)cos(\u03c6) \u2212zbBcos(\u03b8)sin(\u03c6) \u2212zbBsin(\u03b8) 0 (8) Figure 4 shows the forces generated by the thrusters acting on the micro submarine, these are described relative to the body-fixed coordinate system, as: f\u0302 = 0 0 f1 ; f\u0302 = 0 0 f2 ; f\u0302 = f3 0 0 ; f\u0302 = f4 0 0 summarizing and using the notation of [14], it follows that \u03c4 = \u03c4X \u03c4Y \u03c4Z = f3 + f4 0 f1 + f2 (9) and the body-fixed torques generated by the above forces, are defined as: \u03c4 = 4\u2211 i=1 li \u00d7 f\u0302i (10) where li = (lix, liy, liz) is the position vector of the force f\u0302i \u2200 i = 1, .., 4, with respect to the body-fixed reference frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001058_rast.2009.5158295-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001058_rast.2009.5158295-Figure6-1.png", + "caption": "Figure 6 Magnetic Torquers", + "texts": [ + " c) H-Bridge The purpose of the H-bridge is to change the direction of the current that passes through the magnetic torquers producing magnetic moment. By changing the direction of the current, the direction of the magnetic moment is changed, and thus we are able to have control in two directions around each of the axes. The principle of an H-bridge is shown in Figure 5, where the MCU is able to choose the direction of the current through the coils by activating different transistors. d) Magnetic Torquers The magnetic torquers are copper coils that are located at three of the satellite sides as shown in Figure 6 where they enable three axis control. The magnetic torquers are mounted on the outside of the structure and are behind the solar panels (shown as brown lines in Figure 6). e) Microcontroller (MCU) The microcontroller controls all the functions that are defined in Figure 7. Regarding selection of a microcontroller for ADCS, it must be taken into account that it should be able to perform a high level of computations due to the mathematical algorithms implemented in this system. Dependent on the lifetime of the project the radiation resistance of the controller may also be needed to be factored into the selection process. f) Real Time Clock The real time clock is an important part of the ADCS system, as it is required to perform rotations between reference frames, determination of the magnetic field and to determine the sun vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003610_joe.2018.9382-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003610_joe.2018.9382-Figure3-1.png", + "caption": "Fig. 3 Integrated structure of flat voice coil actuator and eddy current sensor", + "texts": [ + " The oscillation voltage passes through the detection, filtering, compensation, amplification, and normalisation, transformed into voltage changes, and then the conversion of displacement to voltage is completed. The eddy current position sensor and the flat VCA are designed as an integrated structure, with the probe fixed on the magnet through the bracket. It has the advantages of high-precision relative position measurement and high-precision force output. The structure is also compact and easy to install. The integrated structure of the eddy sensor and the voice-coil actuator is shown below (see Fig. 3). A mathematical model of output force of the flat VCA is established. In the coordinate system, shown in Fig. 1, Euler angle is adopted to describe the relative attitude parameters of O1X1Y1Z1 and O2X2Y2Z2. Assuming that O1X1Y1Z1 rotates to O2X2Y2Z2 in 3-2-1 order, the attitude matrix between O1X1Y1Z1 and O2X2Y2Z2 could be denoted as (see equation below) where \u03c8 , \u03b8, \u03c6 is attitude angle. The magnetic field intensity direction of the VCA is fixed to the coordinate O1X1Y1Z1, and thus, the coordinate could be described as B1 = Buy; the conductor direction is fixed to the coordinate O2X2Y2Z2, and its coordinates are L2 = lux; the current is I, and the direction is the same as the conductor direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001719_cyberi.2018.8337561-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001719_cyberi.2018.8337561-Figure3-1.png", + "caption": "Fig. 3. Balancing Arm Model.", + "texts": [ + " The first variable is the angle of deviation of the arm from the equilibrium state with respect to the angle of rotation. In the latter case, we can measure the height of the arm end above the base plate. In this case, it is necessary to determine the equilibrium state and also the end state. This is performed by measuring the range of the arm deviation values; actually, a calibration is done so that an appropriate sensor is chosen. We use the first measure option, which is not only easier, but it also provides more precise values. Balancing Arm Model, Fig. 3, was created by commonly available materials, that is steel, stainless steel, aluminum and silon. Basic basement was constructed using steel and creates the biggest part of model. Basement is used to keep a stable position of the arm, and have two pillars on the top. Steel pillars have a fixed height, where on the end are embedded bearing, and on the right pillar, there is a motor holder. Choice of bearing is really important for a correct function of all system, because it represents the major friction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002432_978-3-319-99522-9_8-Figure8.17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002432_978-3-319-99522-9_8-Figure8.17-1.png", + "caption": "Fig. 8.17 Planar 3-RRR parallel robot", + "texts": [ + " Together, the three-degrees-of-freedom manipulator consists of seven moving links, nine revolute joints and three revolute actuators installed on the fixed base. We attach a Cartesian frame Ox0y0z0\u00f0T0\u00de to the fixed base with its origin located at the triangle centre, the z0 axis perpendicular to the base and the x0 axis pointing along the direction C1O. Another mobile reference frame GxGyGzG is attached to the moving platform. The origin of this central coordinate system is located just at the centre G of the moving triangle (Fig. 8.17). In what follows, we consider that the moving platform is initially located at a central configuration, where the platform is not rotated with respect to the fixed base and the mass centre G is at the origin O of the fixed frame. Relative rotation of any body Tk with angle uk;k 1 must be always pointing about the direction of zk axis. First active leg A, for example, consists of a fixed revolute joint A1, a moving crank 1 of length l1, mass m1 and tensor of inertia J\u03021, which has a rotation about z A1 axis with the angle uA 10, the angular velocity xA 10 \u00bc _uA 10 and the angular acceleration eA10 \u00bc \u20acuA 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000752_sice.2008.4655087-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000752_sice.2008.4655087-Figure5-1.png", + "caption": "Fig. 5 Stepping reflex. The robot changes the touchdown angle of the swing leg according to the ankle joint angular velocity of the stance leg.", + "texts": [ + " Backward motion of the swing leg is planned at the hip joint in order for the swing leg to land on ground and move the body forward after landing.The desired trajectory of kick motion is given to the ankle joint of the - 2509 - stance leg to help the body move forward only in the transitional walking.This trajectory was obtained by trial and error. The knee joints of both legs are mechanically locked. The time t is reset to zero at t = T0 + \u2206T , and stance and swing legs are exchanged. A stepping reflex is illustrated in Fig. 5. eq.(3) is used in place of the trajectory of the hip joint of the swing leg in the single leg stance period in Table 1. \u03b8sw h d = \u03c6(t) + ksr \u00d7 (\u03b8\u0307st a c \u2212 \u03b8\u0307st a d) (3) where ksr is the gain of stepping reflex. Using this stepping reflex, a robot can adjust the touchdown angle of the swing leg according to the change of angular momentum of the system. Since this touchdown angle becomes the initial angle of the stance leg in the next step, the robot can adjust the motion of the stance leg moving as an inverted pendulum approximatelywhile utilizing gravity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.25-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.25-1.png", + "caption": "Fig. 11.25 Relation among the relative average sliding velocity V 0, velocity of the road VR and velocity of the tread base VB", + "texts": [ + " If the static friction coefficient ls is the same in all directions, the force equilibrium is expressed by qzls\u00f0 \u00de2\u00bc f 2x \u00fe f 2y ; \u00f011:97\u00de where qz is the parabolic function given by Eq. (11.46). If we assume the relations a 1, s 1 and Cx = Cy = C for simplicity, using Eqs. (11.46), (11.89), (11.90) and (11.97), lh is obtained as11 lh \u00bc l 1 CF 3lsFz ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi tan2 a\u00fe s2 p CF \u00bc Cbl2=2; \u00f011:98\u00de where Fz is the load given by Eq. (11.47). (3) Side force and self-aligning torque in the sliding region The sliding direction in the sliding region is shown in Fig. 11.25, where V 0 is the relative sliding velocity between the tread base and the road, VR is the road velocity, and VB is the velocity of the tread base. Referring to Fig. 11.25, it follows that12 11Note 11.9. 12Note 11.10. s tan h \u00bc tan a 0 h p\u00f0 \u00de: \u00f011:99\u00de The side forces of a tire in the sliding region are given by F00 x \u00bc b Z l lh ldqz cos hdx1 F00 y \u00bc b Z l lh ldqz sin hdx1: \u00f011:100\u00de Assuming that the tread approximately moves on line BC of Fig. 11.24 in the sliding region, the moment is integrated around the z-axis. The self-aligning torque of a tire in the sliding region is given by M00 z \u00bc b Z l lh ldp y0 cos h\u00fe x1 l 2 sin h dx1: \u00f011:101\u00de Here, y0 is the lateral distance of line BC measured from the x-axis: y0 \u00bc x1 l\u00f0 \u00delh tan a lh l \u00fe y0; \u00f011:102\u00de where y0 is given by Eq", + " In the case of braking (s > 0), 4lspm x l 1 x l \u00bc x ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi C2 x s 2 \u00feC2 y sin 2 a q x \u00bc l 1 l 4lspm ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi C2 x s 2 \u00feC2 y sin 2 a q : Assuming Cx = Cy = C and considering pm = 3Fz/(2 lb) and CFa = Cl2b/2, we obtain lh \u00bc l 1 CFa 3lsFz ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 \u00fe sin2 a p ffi l 1 CFa 3lsFz ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 \u00fe tan2 a p : In the case of driving (s < 0), x \u00bc l 1 l 4lspm ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi C2 x s 2 \u00feC2 y \u00f01\u00fe s\u00de2 tan2 a q ; where Cx = Cy = C and, in a manner similar to the case of braking, we obtain lh \u00bc l 1 CFa 3lsFz ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 \u00fe\u00f01\u00fe s\u00de2 tan2 a q ffi l 1 CFa 3lsFz ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 \u00fe tan2 a p : Note 11.10 Eq. (11.99) The following relations are obtained referring to Fig. 11.25. In the case of braking, s \u00bc VR cos a VB VR cos a \u00bc a a\u00feVB ; tan h \u00bc d a ; tan a \u00bc d a\u00feVB ) a a\u00feVB d a \u00bc d a\u00feVB ) s tan h \u00bc tan a: In the case of driving, s \u00bc VR cos a VB VB \u00bc a VB ; tan h tan h0 \u00bc d a ; tan a \u00bc d VB ) a VB d a \u00bc d VB ) s tan h \u00bc tan a: Note 11.11 Eq. (11.104) Using Eq. (11.99), we obtain sin h \u00bc tan a= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 \u00fe tan2 a p \u00bc h tan a cos h \u00bc s= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 \u00fe tan2 a p \u00bc hs: Equation (11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003507_elk-1702-150-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003507_elk-1702-150-Figure1-1.png", + "caption": "Figure 1. SIDM structure.", + "texts": [ + " The experiments show that a prototype of the SIDM and controller can obtain nanoscale positioning accuracy, and the control system can maintain steady operation when the load changes. This paper is organized as follows. In Section 2, the movement principles of an SIDM are analyzed. Section 3 presents the design of the controller, including the NNI and the NNC. Section 4 presents a fabricated prototype of the SIDM and the experimental system. In Section 5, the experimental results are presented under different working conditions. Section 6 provides a discussion and conclusion. In this paper, an SIDM is composed of piezoelectric stack, friction rod, and slider, as shown in Figure 1. When the piezoelectric stack is excited by the electricity, it simultaneously extends and pushes the friction rod, which then moves forward because of the friction between the rod and the slider. The working process of the SIDM can be classified into four moving phases when the SIDM is excited by a trapezoid signal, as illustrated in Figure 2. A\u2013B: refers to the rising edge of the driving voltage, which is the stick stage between the slider and frictional rod. The piezoelectric stack extends slowly, the friction rod is pushed directly forward by the driving friction force, and a small forward distance, \u2206y1 of the slider, is produced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003116_andescon.2018.8564721-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003116_andescon.2018.8564721-Figure1-1.png", + "caption": "Fig. 1. Representaci\u00f3n de los arrollamientos estat\u00f3ricos y rot\u00f3ricos de la m\u00e1quina de inducci\u00f3n.", + "texts": [ + " Finalmente se realiza la simulaci\u00f3n del circuito equivalente del DFIG para diferentes factores de potencia, donde se consideran como variables de entrada el voltaje del estator, el par mec\u00e1nico dado por la turbina e\u00f3lica, la velocidad mec\u00e1nica del eje y el factor de potencia. II. METODOLOG\u00cdA PARA EL AN\u00c1LISIS DEL DFIG EN ESTADO ESTACIONARIO Para obtener el modelo matem\u00e1tico se consideran los siguientes supuestos [8]: - No se considera la saturaci\u00f3n magn\u00e9tica. - La fmm en el entrehierro es de forma sinusoidal. - Se desprecian las p\u00e9rdidas en el hierro. La Fig. 1, muestra la estructura electromagn\u00e9tica de la m\u00e1quina de inducci\u00f3n de entrehierro uniforme con un devanado de tres fases en el estator y tres fases en el rotor, desplazados 120 grados el\u00e9ctricos. Aplicando la ley circuital de voltajes y la ley de inducci\u00f3n a cada una de las seis fases se obtienen tres ecuaciones diferenciales para el estator m\u00e1s tres ecuaciones para el rotor. Las seis ecuaciones determinan el modelo abc, el cual se puede expresar en forma matricial, as\u00ed: = (1) = (2) En donde la letra representa el voltaje por fase, la corriente y representa el flujo concatenado" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003172_ecce.2018.8558445-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003172_ecce.2018.8558445-Figure2-1.png", + "caption": "Figure 2: Affinity laws for pumps and fans in variable torque (VT) mode.", + "texts": [], + "surrounding_texts": [ + "In applications such as heating, ventilating and air conditioning (HVAC), scalar Volts/Hz control mode dominates the industry, due to its simplicity and economic benefits. But, a linear Volts/Hz curve typically provides a voltage higher than necessary, resulting in wasted energy, especially when the reference frequency is significantly lower than the rated frequency. IM core losses are mainly composed of losses caused by eddy current and hysteresis in the iron core, and are typically proportional to the square of the input voltage. Applying unnecessarily high voltage to the motor generates excessive motor core losses in the form of heat and noise. To quantify the effectiveness of energy efficiency optimization, the linear Volts/Hz is defined as a baseline, providing a benchmark for evaluating three different algorithms: (a). Scalar quadratic Volts/Hz curve; (b). Flux optimization based on FOC; (c). Scalar energy optimizing Volts/Hz control." + ] + }, + { + "image_filename": "designv11_92_0002287_s00170-018-2422-y-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002287_s00170-018-2422-y-Figure5-1.png", + "caption": "Fig. 5 Analysis of minimum thickness when \u03b1f is less than \u03b1", + "texts": [ + " L1sin\u03b81 \u00fe 2Risin \u03b81 2 sin \u03b81 2 \u00bc L2sin\u03b82 \u00fe 2Risin \u03b82 2 sin \u03b82 2 Let parameters \u03a6 be defined as sin\u22121 Riffiffiffiffiffiffiffiffiffiffiffiffi L12\u00feRi 2 p , \u03b8 as sin\u22121 Riffiffiffiffiffiffiffiffiffiffiffiffi L22\u00feRi 2 p , T1 as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L12 \u00fe Ri 2 p , and T2 as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L22 \u00fe Ri 2 p , the following equation is obtained: With parameter \u03b2 defined as \u03b2 \u00bc tan\u22121 T2sin \u03d5\u00fe\u03b8\u2212\u0394\u03b1\u00f0 \u00de T1\u00feT2cos \u03d5\u00fe\u03b8\u2212\u0394\u03b1\u00f0 \u00de \u00fe 0 if T 1 \u00fe T2cos \u03d5\u00fe \u03b8\u2212\u0394\u03b1\u00f0 \u00de\u22650 \u03c0 if T1 \u00fe T2cos \u03d5\u00fe \u03b8\u2212\u0394\u03b1\u00f0 \u00de<0 ; angle\u019f1 can be calculated as \u03b81 = \u03d5 \u2212 \u03b2. Using the relations between the parameters described above, \u019f2, t1, t2, and to can then be obtained in turn. (2) \u03b1f is less than \u03b1 When prototyping sheet metals, the angle of post-bending workpiece \u03b1f could be less than the angle of the design model \u03b1. When\u03b1f is less than\u03b1, a similar drawing is made in Fig. 5 to relate the design model and the actual workpiece, wherein the upper boundary of the actual workpiece touches both flanges of the design model. In order to determine the minimum enclosing thickness, the same objective is defined as minimizing the enclosing thickness t + max (t1, t2) and is listed as Minimum thickness = t +min {max(t1, t2)}, where t is the thickness of the design model and t1 and t2 are thickness increases for each flange respectively. Those parameters are related as t1 \u00bc L1 sin \u03b81\u22122Ri sin 2\u03b81 2 \u00f07\u00de t2 \u00bc L2 sin \u03b82\u22122Ri sin 2\u03b82 2 \u00f08\u00de \u03b81 \u00fe \u03b82 \u00bc \u03b1\u2212\u03b1 f \u00bc \u0394\u03b1 \u00f09\u00de When t1 = t2, L1sin\u03b81\u2212 2Ri sin \u03b81 2 sin \u03b81 2 \u00bc L2 sin \u03b82\u2212 2Risin \u03b82 2 sin \u03b82 2 With the definitions \u03a6\u00bc sin\u22121 Riffiffiffiffiffiffiffiffiffiffiffiffi L12\u00feRi 2 p , \u00bc sin\u22121 Riffiffiffiffiffiffiffiffiffiffiffiffi L22\u00feRi 2 p , T1 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L12 \u00fe Ri 2 p and T 2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L22 \u00fe Ri 2 p , we have Wi th \u03b2 de f i n ed a s \u03b2 \u00bc tan\u22121 T2sin \u2212\u0394\u03b1\u2212\u03b8\u2212\u03d5\u00f0 \u00de T1\u00feT2cos \u2212\u0394\u03b1\u2212\u03b8\u2212\u03d5\u00f0 \u00de \u00fe 0 if T1 \u00fe T2cos \u2212\u0394\u03b1\u2212\u03b8\u2212\u03d5\u00f0 \u00de\u22650 \u03c0 if T1 \u00fe T2cos \u2212\u0394\u03b1\u2212\u03b8\u2212\u03d5\u00f0 \u00de<0 ; angle\u019f1 can be calculated as \u03b81 = \u2212\u03d5 \u2212\u03b2, and similarly,\u019f2, t1, t2, and the minimum thickness are then solved in turn", + " The first constraint \u2202Wd \u2201 \u2202Wa leads to the maximum \u019fj, such that \u03b8 j \u00bc sin\u22121 \u0394t\u2212Ri T j \u2212\u03c6*, whereby the boundary of design model barely touches the boundary of the selected workpiece. Meanwhile, considering the second constraint \u2211 2 j\u00bc1 \u03b8 j \u00bc \u0394\u03b1, the upper limit of \u03b8j is obtained as follows \u03b8 jmax \u00bc Min sin\u22121 \u0394t\u2212Ri T j \u2212\u03c6*;\u0394\u03b1 \u00f013\u00de Since the maximum value of \u03b81 corresponds to the minimum value of \u03b82 and vice versa, the limits of both \u03b81 and \u03b82 are determined as \u03b81max \u00bc Min sin\u22121 \u0394t\u2212Ri T1 \u2212\u03c6*;\u0394\u03b1 \u00f014\u00de \u03b82max \u00bc Min sin\u22121 \u0394t\u2212Ri T2 \u2212\u03c6*;\u0394\u03b1 \u00f015\u00de \u03b81min \u00bc \u0394\u03b1\u2212\u03b82max \u00f016\u00de \u03b82min \u00bc \u0394\u03b1\u2212\u03b81max \u00f017\u00de For the second case discussed in this section where \u03b1f is greater than\u03b1 (Fig. 5),\u03c6\u2217 is defined as\u03c6\u2217 = atan2 (Ri, Lj), and the limits of \u03b81 and \u03b82 are determined in a similar way and are listed below \u03b81max \u00bc Min sin\u22121 \u0394t\u00fe Ri T1 \u2212\u03c6*;\u0394\u03b1 \u00f018\u00de \u03b82max \u00bc Min sin\u22121 \u0394t\u00fe Ri T2 \u2212\u03c6*;\u0394\u03b1 \u00f019\u00de \u03b81min \u00bc \u0394\u03b1\u2212\u03b82max \u00f020\u00de \u03b82min \u00bc \u0394\u03b1\u2212\u03b81max \u00f021\u00de Essentially, the machining process after bending in this research belongs to sheet metal machining. It can be seen from Figs. 4 and 5 that when deviation angles \u019f1 and \u019f2 are not equal to zero, the radial depth of machining process linearly changes. Therefore, the cutting force varies as the milling cutter moves along the workpiece contour", + " Phenomena such as thinning and bulging at the deformation zone are neglected. Depending on the relation of final bend angle \u03b1f and design bend angle \u03b1, workpiece length can be estimated as the summation of both flange lengths and arc length at the deformation zone. When \u03b1f is greater than \u03b1 as illustrated in Fig. 4, the workpiece length can be calculated using the following formula. L1 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b81 2 cos\u03b81 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b81 2 \u00fe L2 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b82 2 cos\u03b82 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b82 2 \u00fe \u03c0\u2212\u03b1 f Ri \u00fe t\u2212 top 2 When\u03b1f is less than\u03b1 as illustrated in Fig. 5, the workpiece length can be calculated in a similar way with the following formula L1\u2212Ritan \u03b81 2 cos\u03b81 \u2212 Ritan \u03b81 2 \u00fe L2\u2212Ritan \u03b82 2 cos\u03b82 \u2212 Ritan \u03b82 2 \u00fe \u03c0\u2212\u03b1 f Ri \u00fe top 2 Given the nature of machining process used in this research, the process planning requires a compensation step before toolpath planning. A regular three-axis CNC milling machine is used in this research to remove unnecessary material from the post-bending workpiece. The cutting force will cause workpiece deformation and cutter deflection, both of which are considered as displacements from their nominal positions although they are along opposite directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001794_1.5034601-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001794_1.5034601-Figure5-1.png", + "caption": "FIGURE 5. Outline drawing of the gantry crane KKS-32", + "texts": [ + " It is easy to see that the equations of motion and the boundary conditions are quite analogous to those of (2), (3) and (4), where one should only replace the mass of the entire crane m by the mass of the load trolley itself m~ , the distance and the time of motion 1S and 1 ~T by the respective values 2S and 2 ~T of the second subtask. Applying the Pontryagin maximum principle as in the previous case, we obtain the control force ,~~, ~ ~ ,sincos 122726542 mmM lm gMtCtCtCCtF while the generalized Gauss principle gives .~~~~~ 3 7 2 6542 tCtCtCCtF EXAMPLE OF A NUMERICAL SOLUTION Consider for example the operation of a real crane KKS-32 (Fig. 5). The mass of the entire crane is m = 94 tons, the mass of the load trolley is m~ = 7 tons, and the maximum load capacity is 32 tons. A container of a mass 1m = 25 tons is to be transferred to the point C (Fig. 2) so that 1S = 218 meters, and 2S = 46 meters. The length of the cable holding the load is l = 11 meters. An important limitation is the maximum speed of the crane and of the load trolley limited by the value 0.57 m/s, hence, the time of motion 383 57.0 ~ 1 1 S T seconds, and 81 57.0 ~ 2 2 S T seconds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001614_access.2018.2811759-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001614_access.2018.2811759-Figure1-1.png", + "caption": "FIGURE 1. Constant disturbance in the orbital plane.", + "texts": [ + " Because a disturbance in pitch can be offset by a constant bias in the pitch axis according to (19), we assume that 1H\u0307 o s (2) = 0 in (18), and hence the dynamics of the space station in the pitch axis can be rewritten as( \u2212ty \u2212 3\u03c92 oI12 ) \u03d5 + [ 3\u03c92 o (I33 \u2212 I11)+ tx ] \u03d1 + 3\u03c92 oI23\u03c8 + T o dy = 3\u03c92 oI13 (26) Ignoring the impact of the coupling effect between roll/ yaw and pitch, we decouple the pitch axis from the yaw/roll axes. The TEA in pitch can be estimated using the following equation: \u03b80y = ( T oy0 \u2212 3\u03c92 oI13 ) / [ 3\u03c92 o(I11 \u2212 I33)\u2212 tx ] (27) The following analysis will show that the attitude error in yaw can be held to zero by the roll axis undergoing a sinusoidal motion at the orbital frequency. The estimated error in pitch caused by the decoupling is very small. To make the explanation more visual, the constant disturbance in the orbital plane is illustrated in Fig. 1. The vector \u2212\u2192 AB represents the orientation of the constant component of the disturbance torque T id in the orbital plane, and \u03b1 is the angle between T idxz0 and ox i. The inertial and LVLH frames overlap at the initial time. The angle between xo and xi is\u03c9ot at time t , and the transformation matrix from the LVLH frame to the inertial frame is defined as C i o = cos\u03c9ot 0 \u2212 sin\u03c9ot 0 1 0 sin\u03c9ot 0 cos\u03c9ot (28) To obtain the constant component of T id in the inertial frame,T id0 = [ T ix0 T i y0 T i x0 ]T , we need to express the orbital frequency component of T0 d in the inertial frame and retain the constant term: T id0 = 0.5(b\u2212 c) T oy0 0.5(a+ d) (29) This means \u2223\u2223T idxz0\u2223\u2223 = 0.5 \u221a (b\u2212 c)2 + (a+ d)2 is deter- mined by the phases of Tod at the orbital frequency, and \u03b1 = arccos [ 0.5(b\u2212 c)/ \u2223\u2223\u2223T idxz0\u2223\u2223\u2223] (30) 15078 VOLUME 6, 2018 The disturbance in the orbital plane shown in Fig. 1 must be offset by the roll/yaw motion. However, the aerodynamic torque is very small, the yawmotion can produce very limited torques according to (19), and we utilize the roll axis motion to produce a control torque in the orbital plane and keep the attitude stable in yaw. When the space station operates in the orbital arc A\u0302B, the roll should produce a control torque in the opposite direction of the roll axis to offset the disturbance in the orbital plane. Similarly, when the space station is operating on the orbital arc B\u0302A, the roll should produce a control torque in the positive direction of ooxo to offset the disturbance in the orbital plane", + "01/t5s3 ax2 = (300\u03d52 + 200ts3\u03d5\u03072 \u2212 2) ts3 bx2 = (1\u2212 400\u03d52 \u2212 300ts3\u03d5\u03072) t2s3 ex2 = 100t5s3\u03d5\u03072{ \u03d51 = \u2212kx cos (\u03c90t2 \u2212 \u03b1) sign (I33 \u2212 I22) \u03d52 = \u2212kx cos (\u03c90t3 \u2212 \u03b1) sign (I33 \u2212 I22) (39) kx is the maximum steady state attitude, which can be estimated using the initial CMGs angular momentum and attitude related coefficient: kx = \u2223\u2223\u2223\u03c9ohiucxz/ (2\u03c0n)+ T ixzd0\u2223\u2223\u2223 / [1.5\u03c92 o (I33 \u2212 I22) ] (40) Because the angular momentum of CMGs in the orbital plane, hiucxz, is generated using the constant disturbance in the orbital plane, T ixzd0, it is reasonable to assume that hiucxz is parallel to T ixzd0, which means the phase angle \u03b1 in (39) is the same as the phase angle shown in Fig. 1. NOTE 1: The attitude path planned in this part is based on the dynamic characteristics and CMGs angular momentum that needs to be unloaded. The effect of unloading depends on the accuracies of the models, especially the disturbance model, and it is therefore difficult to unload the CMGs angular momentum to precisely zero. IV. CONTROLLER DESIGN A. TORQUE EQUILIBRIUM ATTITUDE TRACKING Because of the large angles in attitude manoeuvres, the linearized model cannot be used to design the TEA tracking controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001226_kikaic.74.2870-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001226_kikaic.74.2870-Figure1-1.png", + "caption": "Fig. 1 Ball hit on its upper left part with inclined cue", + "texts": [], + "surrounding_texts": [ + "2870\n\u65e5\u672c\u6a5f\u68b0\u5b66 \u4f1a\u8ad6 \u6587\u96c6(C\u7de8)\n74\u5dfb748\u53f7(2008-12)\n\u8ad6 \u6587No.08-0501\n\u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b \u304a \u3051\u308b \u30ad \u30e5\u30fc\u306e\u885d \u7a81 \u7279\u6027 \u8a55 \u4fa1*\n(\u7403\u306e\u8ecc\u9053\u3092\u30ab\u30fc\u30d6\u3055\u305b\u308b\u5834\u5408)\n\u5cf6 \u6751 \u771f \u4ecb*1, \u9808 \u8cc0 \u4e00 \u535a*2 \u6c5f \u6fa4 \u826f \u5b5d*3, \u9752 \u6728 \u7e41*4\nNumerical Evaluation of Impact Characteristics of Cue in Billiards\n(Case Where Ball Trajectory is Curved)\nShinsuke SHIMAMURA*5, Kazuhiro SUGA, Yoshitaka EZAWA and Shigeru AOKI\n*5 Graduate School of Engineering, Toyo University, 2100 Kujirai, Kawagoe-shi, Saitama, 350-8585 Japan\nThe numerical evaluation of impact characteristics of a cue in billiards is performed for the case where the ball trajectory is curved by hitting a ball on its upper left or right part with an inclined cue. The effective numerical method developed in the previous papers, in which a cue and a ball are assumed to be isotropically elastic and rigid, respectively, is extended for this case. The extended method is verified with an experiment using a high speed camera. The result obtained with the extended method also shows that the curvature of the ball trajectory is large for the impact on its almost exactly left part with a slightly inclined cue, and this result agrees with an empirical one. It is found from a numerical simulation that if the Young's modulus or mass density of a shaft of a cue is large, the ball trajectory immediately after the impact is deviated largely and then is curved with a large curvature. Because the extended method can evaluate quantitatively this kind of effects of material properties, it is useful for a design of a cue.\nKey Words : Impact, Cue, Billiards, Curved Ball Trajectory, Efficient Numerical Method , Simulation\n1. \u7dd2 \u8a00\n\u30d3\u30ea\u30e4\u30fc \u30c9\u306f,\u30c6 \u30cb\u30b9\u3084\u30b4\u30eb \u30d5\u3068\u3068\u3082\u306b\u4eba\u6c17 \u304c\u9ad8 \u304f\n\u4e16 \u754c\u4e2d\u3067\u89aa \u3057\u307e\u308c\u3066 \u3044\u308b\u304c,\u30ad \u30e5\u30fc\u306e\u8a2d\u8a08\u306b\u95a2\u3059 \u308b\u7814\n\u7a76 \u306f\u5c11\u306a\u3044.\u30ad \u30e5\u30fc\u306f\u5148\u7aef\u304b \u3089\u30bf \u30c3\u30d7,\u5148 \u89d2(\u3055 \u304d\u3064 \u306e)\u304a \u3088\u3073\u30b7\u30e3\u30d5 \u30c8\u3067\u69cb\u6210\u3055\u308c,\u305d \u308c\u305e \u308c\u901a\u5e38,\u9769, \u30d7\u30e9\u30b9\u30c1\u30c3\u30af\u304a \u3088\u3073\u6728\u6750\u3067\u4f5c\u6210\u3055\u308c\u308b.\u30b7 \u30e3\u30d5 \u30c8\u306f\u524d\n\u534a\u5206 \u3068\u5f8c \u534a\u5206 \u306b\u5206\u304b\u308c,\u4e2d \u592e \u3092\u91d1\u5c5e\u88fd\u306e\u306d \u3058\u3067\u9023 \u7d50\u3055 \u308c\u308b \u3053\u3068\u304c\u591a \u3044.\n\u307e\u305f,\u30ad \u30e5\u30fc\u306f\u7af6\u6280\u306e\u5c40\u9762\u306b\u5fdc \u3058\u3066(1)\u7403 \u306e\u4e2d\u592e\u90e8 \u3092\u649e \u304f\u57fa \u672c\u7684\u306a\u649e \u304d\u65b9\u306e\u4ed6\u306b,(2)\u4e0a \u90e8\u307e\u305f\u306f\u4e0b\u90e8\u3092\n\u649e\u3044\u3066\u7403\u306b\u6b63\u56de\u8ee2 \u307e\u305f \u306f\u9006\u540c\u8ee2 \u3092\u4e0e\u3048\u308b\u649e\u304d\u65b9,(3)\n\u6a2a\u90e8\u3092\u649e\u3044\u3066\u7403\u306b\u30b9\u30d4\u30f3\u3092\u4e0e\u3048\u308b\u649e\u304d\u65b9,(4)\u659c \u3081\u4e0a \u90e8\u3092\u659c\u3081\u4e0b\u65b9\u306b\u649e\u3044\u3066\u7403\u306e\u8ecc\u9053\u3092\u30ab\u30fc\u30d6\u3055\u305b\u308b\u649e\u304d\u65b9 \u306a\u3069\u69d8\u3005\u306a\u649e\u304d\u65b9\u304c\u8981\u6c42\u3055\u308c\u308b.\n\u30ad\u30e5\u30fc\u3092\u5408\u7406\u7684\u306b\u8a2d\u8a08\u3059\u308b\u305f\u3081\u306b\u306f,\u30ad \u30e5\u30fc\u5404\u90e8\u306e \u5f62\u72b6,\u5bf8 \u6cd5\u304a\u3088\u3073\u6750\u8cea\u304c\u3053\u306e\u3088\u3046\u306a\u69d8\u3005\u306a\u649e\u304d\u65b9\u306b\u3088 \u308b\u885d\u6483\u529b,\u7403 \u306e\u904b\u52d5\u306a\u3069\u306e\u885d\u7a81\u7279\u6027\u306b\u4e0e\u3048\u308b\u5f71\u97ff\u3092\u5b9a \u91cf\u7684\u306b\u8a55\u4fa1\u3059\u308b\u5fc5\u8981\u304c\u3042\u308b.\u7b46 \u8005\u3089\u306f\u524d\u5831(1)\uff5e(3)\u3067, \u7403\u3092\u525b\u4f53,\u30ad \u30e5\u30fc\u3092\u5f3e\u6027\u4f53\u3068\u4eee\u5b9a\u3057,\u305d \u308c\u305e\u308c\u3092\u6709\u9650 \u5dee\u5206\u6cd5\u304a\u3088\u3073\u6709\u9650\u8981\u7d20\u6cd5\u3067\u89e3\u6790\u3057,\u4e21 \u8005\u3092\u9023\u6210\u3059\u308b\u3053 \u3068\u306b\u3088\u308a,\u4e0a\u8a18(1)\uff5e(3)\u306e \u649e\u304d\u65b9\u306b\u5bfe\u3059\u308b\u885d\u7a81\u7279 \u6027\u3092\u8a55\u4fa1\u3059\u308b\u7c21\u4fbf\u306a\u65b9\u6cd5(\u52d5 \u7684\u63a5\u89e6\u6709\u9650\u8981\u7d20\u6cd5\u306b\u6bd4\u3079 \u3066\u8a08\u7b97\u6642\u9593\u304c\u7d041/30)\u3092 \u958b\u767a\u3057\u305f.\n\u672c\u7814\u7a76\u3067\u306f\u524d\u5831\u3067\u958b\u767a\u3057\u305f\u65b9\u6cd5\u3092,\u4e0a\u8a18(4)\u306e \u649e\u304d\n\u65b9\u306b\u3082\u9069\u7528\u51fa\u6765\u308b\u3088\u3046\u306b\u62e1\u5f35\u3057,\u30b7 \u30e3\u30d5\u30c8\u306e\u6750\u8cea\u304c\u885d \u7a81\u7279\u6027\u306b\u4e0e\u3048\u308b\u5f71\u97ff\u3092\u660e\u3089\u304b\u306b\u3059\u308b.\u524d \u5831\u3068\u540c\u69d8\u306b, \u533a\u5225\u306e\u5fc5\u8981\u304c\u3042\u308b\u5834\u5408\u306b\u306f,\u53f0 \u306b\u5782\u76f4\u306a\u8ef8\u307e\u308f\u308a\u306e\u7403 \u306e\u56de\u8ee2\u3092\u30b9\u30d4\u30f3,\u305d \u306e\u4ed6\u306e\u8ef8\u307e\u308f\u308a\u306e\u56de\u8ee2\u3092\u8ee2\u304c\u308a\u3068\n\u547c\u3076\u3053\u3068\u306b\u3059\u308b.\n* \u539f\u7a3f\u53d7\u4ed82008\u5e746\u67082\u65e5. *1\n\u6771\u6d0b \u5927\u5b66 \u5927\u5b66 \u9662\u6a5f \u80fd \u30b7 \u30b9\u30c6 \u30e0\u5c02\u653b(\u3013350-8585\u5ddd \u8d8a \u5e02\u9be8 \u4e95\n2100). *2\n\u6771\u4eac\u7406 \u79d1\u5927 \u5b66\u7406\u5de5 \u5b66\u90e8(\u3013278-8510\u91ce \u7530\u5e02 \u5c71\u5d0e2641). *3 \u6b63 \u54e1,\u6771 \u6d0b\u5927 \u5b66 \u30b3 \u30f3 \u30d4\u30e5 \u30c6 \u30fc \u30b7 \u30e7\u30ca \u30eb \u5de5 \u5b66 \u79d1(\u3013350-8585\n\u5ddd\u8d8a \u5e02\u9be8\u4e952100). *4\n\u6771\u6d0b \u5927\u5b66 \u30b3 \u30f3\u30d4\u30e5\u30c6 \u30fc\u30b7 \u30e7\u30ca\u30eb\u5de5\u5b66 \u79d1.\nE-mail:ezawa@toyonet.toyo.ac.jp", + "\u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b\u304a \u3051 \u308b\u30ad\u30e5\u30fc \u306e\u885d\u7a81 \u7279\u6027 \u8a55 \u4fa1 2871\n2. \u89e3 \u6790\u65b9\u6cd5\u306e\u6982\u8981\n\u56f31\u306b \u793a\u3059\u3088 \u3046\u306b,\u7403 \u306e\u659c\u3081\u4e0a\u90e8 \u3092\u659c\u3081\u4e0b\u65b9\u306b\u649e \u3044\n\u3066,\u7403 \u306e\u8ecc\u9053\u3092\u30ab\u30fc\u30d6 \u3055\u305b\u308b\u5834\u5408\u3092\u8003\u3048\u308b.\u305f \u3060 \u3057,\n\u7403 \u3068\u5e8a\u306e\u63a5\u89e6\u529b\u306b\u3088\u308b\u4e21\u8005\u306e\u5909\u5f62\u306f\u7121\u8996\u3067\u304d\u308b\u307b \u3069\u5c0f \u3055\u3044\u3068\u4eee\u5b9a\u3059\u308b.\u524d \u5831(1)\uff5e(3)\u3068 \u540c\u69d8 \u306b,\u30ad \u30e5\u30fc\u5404\u90e8 \u3092\n\u7b49\u65b9\u5f3e\u6027\u4f53,\u7403 \u3068\u53f0\u3092\u525b\u4f53,\u30ad \u30e5\u30fc \u3068\u7403\u306e\u554f\u306e\u885d \u6483\u529b \u3092\u96c6\u4e2d\u8377\u91cd \u3068\u4eee\u5b9a\u3059\u308b.\n\u307e\u305a,\u885d \u6483\u529b \u3092\u65e2\u77e5 \u3068\u4eee\u5b9a \u3057\u3066,\u7403 \u306e\u904b\u52d5\u65b9\u7a0b \u5f0f\u3092\n\u6709\u9650\u5dee \u5206\u6cd5\u3067\u89e3\u304d,\u885d \u6483\u70b9\u306b\u304a\u3051\u308b\u7403\u306e\u5909\u4f4d 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\u3046\u306b,\u925b \u76f4\u4e0b\u65b9 \u3092y\u65b9 \u5411\u3068\u3057,\u30ad \u30e5\u30fc \u306b\u5782\u76f4\u306a\u65b9\u5411 \u3092z\u65b9 \u5411 \u3068\u3059 \u308b\u56fa \u5b9a\u76f4\u89d2\u5ea7\u6a19\u7cfbO-xy,z\u3092 \u63a1\u7528\u3059 \u308b.\u3053 \u306e\u5ea7\u6a19\u7cfb \u3092,\u539f \u70b9\u304c\u904b\u52d5\u3059 \u308b\u7403\u306e \u4e2d\u5fc3 \u306b\u306a \u308b\u3088\u3046\u306b\u5e73\u884c\u79fb\u52d5 \u3057\u305f\u5ea7\u6a19 \u7cfbO'-x'zy',z'\u3082\u7528\u3044\u308b.\u56f31\u306e \u89d2\u5ea6 \u03b1,\u03b2\u3067\u8868 \u3055\u308c \u308b\u7403 \u8868\u9762 \u306e\u70b9 \u3092\u649e \u304f\u5834\u5408 \u3092\u8003 \u3048\u308b.\u7403\u306e\u904b\u52d5\u65b9\u7a0b\u5f0f\u306f\u5f0f(1) \uff5e(6)\u3067 \u4e0e\u3048 \u3089\u308c\u308b.\n\u4e26\u9032\u904b\u52d5:\n(1)\n(2)\n(3)\n\u56de\u8ee2\u904b\u52d5:\n(4)\n(5)\n(6)\n\u305f\u3060 \u3057,\u03b1(t)\u304a\u3088\u3073 \u03b2(t)\u306f\u885d\u6483\u70b9 \u306e\u4f4d\u7f6e \u3092\u8868\u3059\u89d2\u5ea6\u3067\n\u3042 \u308a,\u56f31\u306e \u77e2\u5370\u306e\u65b9\u5411\u3092\u6b63\u3068\u3059 \u308b.b(t)\u306f \u7403 \u4e2d\u5fc3 \u306e \u5909\u4f4d\u30d9\u30af \u30c8\u30eb(\u7d30 \u5b57\u306f\u305d\u306e\u5927 \u304d\u3055,\u4ee5 \u4e0b \u540c\u69d8),t \u304a\u3088\u3073 \u30c9\u30c3 \u30c8(.)\u306f \u305d\u308c\u305e\u308c\u6642\u9593\u304a\u3088\u3073\u6642\u9593\u5fae\u5206, \u4e0b\u6dfb\u3048\u5b57x,y\u304a \u3088\u3073Z\u306f \u305d\u308c\u305e\u308cx,y\u304a \u3088\u3073Z\u65b9 \u5411 \u6210\u5206,f(t)\u306f \u885d\u6483 \u529b\u30d9 \u30af \u30c8\u30eb,Fr(t)\u306f \u6469\u64e6 \u529b\u30d9\u30af \u30c8\u30eb, Mr\u304a \u3088\u3073Ms\u306f \u305d\u308c\u305e\u308c\u8ee2\u304c \u308a\u304a\u3088\u3073\u30b9\u30d4\u30f3\u306b\u5bfe\u3059 \u308b \u6469\u64e6\u30e2\u30fc \u30e1\u30f3 \u30c8,\u03b8(t)\u306f\u7403\u306e\u56de\u8ee2\u89d2\u3092\u8868\u308f\u3059 \u30d9\u30af \u30c8\u30eb, M\u304a \u3088\u30730\u306f 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\u78ba \u306b \u306f \u901f \u5ea6/R)\u3092\n\u03b1\u304a \u3088\u3073 \u03b2\u3067 \u8868 \u308f \u3057,\u53f3 \u8fba \u306f \u03b8x,\u03b8y\u304a \u3088\u3073 \u03b8z\u3067\u8868 \u308f \u3057\n\u3066\u3044 \u308b(\u4ed8 \u9332 \u53c2 \u7167),\u5f0f(15)\u306f,a\u3092 \u03b8x\u3068 \u03b8z\u3067\u8868 \u308f\n\u3057\u3066 \u3044 \u308b\u306e \u3067,\u3053 \u308c \u3092\u5f0f(14)\u307e \u305f \u306f(16)\u306b \u4ee3\u5165 \u3059\n\u308b \u3053 \u3068\u306b \u3088 \u308a,\u03b2 \u3092 \u03b8x,\u306e \u304a \u3088\u3073 \u03b8z\u3067\u8868 \u308f\u3059 \u3053 \u3068\u304c \u51fa\n\u6765 \u308b.\n3\u30fb3 \u89e3 \u6cd5 \u3042 \u308b\u6642 \u523bt\u306b \u304a \u3051 \u308bfx(t),fy(t)\u304a \u3088\u3073\n\u53cc')\u304c \u4e0e \u3048 \u3089\u308c \u305f\u5834 \u5408\u3092 \u8003 \u3048 \u308b.\u307e \u305a,S\u22600\u3068 \u4eee \u5b9a\u3059\n\u308b.Frx(t)\u3068Frz(t)\u306f \u5f0f(17)\u3068 (18)\u304b \u3089 \u03b4(t)\u306b\u3088 \u308a\u8868\n\u3055\u308c \u308b \u306e\u3067,\u6c42 \u3081 \u308b\u3079 \u304d\u5909\u6570 \u306fbx(t),by(t),bz(t),\u03b8x(t),\n\u03b8y(t),\u03b8z(t),s(t),sx(t),sz(t),\u03c1(t),\u03b4(t),\u03b3(t),\u03b1(t)\u304a \u3088\n\u3073 \u03b2(t)\u306e14\u500b \u3067 \u3042 \u308b.\nS\u22600\u306e \u5834 \u5408\n(17)\n(18)\n\u3053\u3053\u3067,\u03bc \u306f\u6469\u64e6 \u4fc2\u6570 \u3092\u8868 \u3057,\u7c21 \u5358 \u306e\u305f \u3081\u306b,\u52d5 \u6469\n\u64e6 \u4fc2\u6570 \u3068\u9759\u6469\u64e6 \u4fc2\u6570 \u3092\u533a\u5225\u305b\u305a\u306b0\u5b9a \u5024 \u3068\u3057\u3066\u3044 \u308b.g \u306f\u91cd\u529b\u52a0 \u901f\u5ea6 \u3067\u3042\u308b.\n14\u500b \u306e\u5909\u6570 \u306e\u5185,\u03b1(t)\u3068\u03b2(t)\u306f\u5f0f(15)\u3068(14)(\u307e\n\u305f\u306f(16))\u304b \u3089 \u03b8x(t),\u03b8y(t),\u304a\u3088\u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b. by(t)\u306f\u65e2\u77e5(by(t)=0)\u3067 \u3042\u308b.S(t),\u03c1(t),\u03b4(t)\u304a\u3088\u3073y(t)\n\u306f\u5f0f(10)\uff5e(13)\u304b \u3089sx(t),sz(t),\u03b8x(t),\u03b8y(t)\u304a\u3088 \u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b.sx(t)\u3068sz(t)\u306f \u5f0f(7)\u3068(9) \u304b \u3089bx(t),bz(t),\u03b8x(t)\u304a\u3088\u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b.\u3057 \u305f\u304c \u3063\u3066,5\u500b \u306e\u672a\u77e5\u6570bx(t),bz(t),\u03b8x(t),\u03b8y(t)\u304a\u3088\u3073 \u03b8z(t) \u306f\u5f0f(1),(3)\uff5e(6)\u306e5\u500b \u306e\u65b9\u7a0b\u5f0f \u304b \u3089\u89e3 \u304f\u3053\u3068 \u304c\u51fa\u6765 \u308b.\u521d \u671f\u6761\u4ef6 \u306f\u3053\u308c \u3089\u306e5\u500b \u306e\u5909\u6570\u304a \u3088\u3073\u305d\u306e\n\u6642\u9593\u5fae \u5206=0\u3067 \u3042\u308b.\u3053 \u308c \u3089\u306e\u5fae\u5206\u65b9\u7a0b\u5f0f \u306f\u6709\u9650\u5dee\u5206 \u6cd5\n(\u4f8b\u3048\u3070,Newmark\u306e \u03b2\u6cd5(\u03b2=1/6)(5))\u306b \u3088 \u308a\u89e3 \u304f\u3053\n\u3068\u304c\u51fa\u6765\u308b.\n\u3082\u3057,\u89e3 \u3044\u305f\u7d50\u679c\u6700\u521d\u306e\u4eee\u5b9a\u304c\u6210 \u7acb\u305b\u305a,S=0\u306b\n\u306a\u3063\u305f\u5834 \u5408\u306f,\u524d \u5831G)\u3068 \u540c\u69d8 \u306b\u5f0f(17)\u3068(18)\u306f \u6210 \u7acb \u3057\u306a\u3044\u305f\u3081,Frx(t)\u3068Frz(t)\u304c\u672a\u77e5\u6570 \u3068\u306a\u308b\u304c,\u7897)\u3068\n\u7897)\u304c0\u3067 \u3042\u308b\u304b \u3089,\u5f0f(7)\u3068(9)\u306b \u3088\u308abx(t)\u3068bz(t) \u306f \u03b8x(t)\u3068\u03b8z(t)\u3067\u8868 \u308f\u3055\u308c \u308b.\u3057 \u305f\u304c \u3063\u3066\u672a \u77e5\u6570 \u306e\u6570 \u3068 \u65b9\u7a0b\u5f0f \u306e\u6570\u304c0\u81f4 \u3057\u89e3 \u304f\u3053\u3068\u304c\u51fa\u6765 \u308b.\n\u306a\u304a,\u7d0410-3\u79d2 \u7a0b\u5ea6\u306e\u30ad\u30e5\u30fc \u3068\u7403 \u306e\u63a5\u89e6\u6642\u9593\u5185\u3067\na(t)=0\u3068 \u4eee\u5b9a(\u524d \u5831(1)\u306b \u3088 \u308a\u7403\u306e \u4e2d\u592e\u3092\u649e \u3044\u305f\u5834\u5408 \u306b \u306f\u63a5 \u89e6\u6642\u9593\u5185\u3067\u306f\u6b86 \u3069\u3059\u3079 \u308a\u306e\u307f\u3067\u56de\u8ee2\u306f\u7121\u8996\u3067 \u304d\n\u308b)\u3059 \u308b\u3068,\u4e0a \u8ff0\u306e\u65b9\u6cd5 \u306f\u524d\u5831(3)\u306e \u65b9\u6cd5 \u306b\u5e30\u7740\u3059 \u308b.\n4. \u4efb\u610f\u65b9 \u5411\u8377\u91cd \u3092\u53d7\u3051\u308b\u30ad\u30e5\u30fc\u306e\u5909\u4f4d\n\u56f31\u306b \u793a\u3059 \u3088\u3046\u306b,\u30ad \u30e5\u30fc \u3092\u659c \u3081\u4e0b\u65b9\u306b\u50be \u3051\u3066,(\u30ad\n\u30e5\u30fc\u306e\u8ef8\u65b9\u5411\u306b)\u7403 \u306e\u659c\u3081\u4e0a\u90e8\u3092\u649e \u304f\u5834\u5408 \u306b\u306f,\u30ad \u30e5 \u30fc\u306b\u306fx\n,y\u304a \u3088\u3073Z\u65b9 \u5411\u306e\u8377 \u91cdfx(t),fy(t)\u304a\u3088\u3073fz(t)\u304c\n\u4f5c\u7528\u3059\u308b.\u3053 \u306e\u5834\u5408\u306e\u885d\u6483\u70b9\u306b\u304a \u3051\u308b\u30ad\u30e5\u30fc\u306e\u5909\u4f4d \u3092 Cx(t),cy(t)\u304a\u3088\u3073cz(t)\u3068\u3057,\u3053 \u308c \u3089\u3092\u52d5\u7684\u5f3e\u6027 \u6709\u9650\u8981\u7d20 \u6cd5\u306b \u3088\u308a\u6c42 \u3081\u308b.\u89e3 \u6790\u306b\u306f,Nastran(6)\u3092 \u7528\u3044\u305f.\n\u88681\u306e \u89e3\u67901(Normal)\u6b04 \u306e\u6750\u6599 \u5b9a\u6570 \u3092\u6301\u3064,\u9577 \u3055 1.48m,\u5148 \u7aef\u90e8\u76f4\u5f840.013m,\u5f8c \u7aef\u90e8\u76f4\u5f840.032m\u306e \u30ad\u30e5 \u30fc\u3092\u8003 \u3048\u308b.\u56f33(a)\u306b \u30ad\u30e5\u30fc\u306e\u30e1 \u30c3\u30b7\u30e5\u5206\u5272\u56f3 \u3092\u793a\n\u3059.\u56f33(b)\u306b \u5185\u90e8\u306e\u69cb\u9020 \u3092\u793a\u3059\u305f\u3081\u306b\u65ad\u9762\u56f3\u3092\u793a\u3059. 8\u7bc0 \u70b9\u306e6\u9762 \u4f53\u8981\u7d20\u307e\u305f\u306f6\u7bc0 \u70b9\u306e5\u9762 \u4f53\u8981\u7d20(\u5168 \u8981 \u7d20\u65706363\u500b,\u5168 \u7bc0\u70b9\u65706184\u500b)\u3092 \u7528 \u3044\u3066\u5206\u5272 \u3057\u3066 \u3044\u308b.\u306a \u304a,\u306f \u308a\u67f1\u8981\u7d20 \u3092\u7528\u3044 \u308b\u3068\u3088 \u308a\u7c21\u4fbf \u306a\u65b9\u6cd5\u304c\n\u5f97 \u3089\u308c \u308b\u3068\u601d\u308f\u308c \u308b\u304c,\u4eca \u5f8c\u306e\u8ab2\u984c \u3068\u3057\u305f\u3044.\n\u885d\u6483\u529b \u3092\u96c6 \u4e2d\u8377\u91cd \u3068\u3057\u305f\u305f\u3081\u306b\u885d \u6483\u70b9 \u306e\u5909\u4f4d\u306b\u8aa4\n\u5dee\u304c\u751f \u3058\u308b\u304a\u305d\u308c \u304c\u3042\u308b\u306e\u3067,\u3053 \u308c \u3092\u5c0f \u3055\u304f\u3059 \u308b\u305f\u3081 \u306b,\u30d8 \u30eb \u30c4\u306e\u63a5\u89e6\u7406\u8ad6(7)\u3067\u4e88\u60f3 \u3055\u308c \u308b\u63a5\u89e6\u534a\u5f84\u306b\u5408 \u308f \u305b\u3066\u885d\u6483\u70b9\u4ed8\u8fd1 \u306e\u8981\u7d20\u306e\u5927\u304d \u3055\u3092\u6c7a\u5b9a \u3057\u305f.\u3059 \u306a\u308f\u3061 \u4e88\u60f3\u6700\u5927\u63a5\u89e6 \u534a\u5f84(\u304a \u3088\u305d2mm)\u306e1/2\u500d \u306e1mm\u3068" + ] + }, + { + "image_filename": "designv11_92_0002810_tastj.16.635-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002810_tastj.16.635-Figure6-1.png", + "caption": "Fig. 6. Photograph of LRECM landing gear.", + "texts": [ + " To validate the effectiveness of the 3D-LRECM, a free-fall landing experiments on a 15-degree slope were conducted. The effectiveness of the 3D-LRECM is confirmed through the comparison to a Spring-Ratchet Mechanism (SRM). This comparison investigates the effect of the tipping over suppression by the counter torque. The SRM locks the kinetic energy absorbed by the spring into the elastic energy of the spring. The SRM landing gear is reproduced by locking the reaction wheel of the LRECM landing gear to prevent the generation of the counter torque. Figure 6 shows a fabricated LRECM landing gear. Similar to Fig. 2, it is composed of a reaction wheel, a spring, latches, telescopic gears, and a footpad. Four such LRECM landing gears were used to construct a 3D-LRECM lander. Figure 7 displays a photograph of the fabricated 3D-LRECM lander and depicts the coordinate system used in the experiments. As an experimental system, a microcomputer was mounted on the center of gravity of the experimental set up, and the accelerations and angular velocities were measured by built-in sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000787_mfi.2008.4648048-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000787_mfi.2008.4648048-Figure5-1.png", + "caption": "Fig. 5. Walking phases on the frontal plane", + "texts": [ + " The most common DCT definition of a 1-D sequence of length N is given in Eq. (5). Similarly, the IDCT is given in Eq. (6). 1 0 (2 1) [ ] ( ) ( ) cos[ ] 2 1 for 0 ( ) 2 for 1, 2, , 1 N n n k X k C k X n N k N C k k N N \u03c0\u2212 = + = = = = \u2212 (5) 1 0 (2 1) ( ) ( ) [ ] cos[ ] 2 N k n k x n C k X k N \u03c0\u2212 = + = (6) Most of the torque signal information tends to be concentrated in low-frequency components of the DCT. Here it is assumed that the joint angles of the left and right legs projected on the frontal plane are the same as frontal\u03b8 as shown in Fig.5 during a walking period. This can be calculated by estimating the ZMP of the biped robot which is the point on the ground about which the sum of all the moments of the active forces equals zero. The position ( zmpx , zmpy ) of the ZMP is represented with respect to the reference frame by Eq. (8).[6] 1 1 1 1 ( ) ( ) n n n i i i i i i iy iy i i i ZMP n i i i m z g x m x z I x m z g = = = = + \u2212 \u2212 \u03a9 = + (8) 1 1 1 1 ( ) ( ) n n n i i i i i i ix ix i i i ZMP n i i i m z g y m y z I y m z g = = = = + \u2212 \u2212 \u03a9 = + where i m is the mass of link i, ix I and iy I are the inertial component, ix \u03a9 and iy \u03a9 are absolute angular acceleration components, g is the gravitational acceleration, ( , , )i i ix y z is the coordinate of the mass center of link i in an absolute Cartesian coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003373_imece2018-86310-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003373_imece2018-86310-Figure3-1.png", + "caption": "Fig. 3: The schematic of the dog-bone shape sample used to determine the mechanical properties of 3D printed steel structure across its thickness, length and height. The gage section measures 200 microns by 200 microns by 1000 microns.", + "texts": [ + " The arc welding process was performed without shield gas. To illustrate the process, a picture of the 3D printer and its schematic are shown in Fig. 1. The movement of the torch is depicted in Fig. 2 along with the directions including tool path, growth and thickness. 2 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 01/21/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use Once solidified, the sample was removed from the base plate and machined with a CNC to produce bow-tie samples shown in Fig. 3. A 1mm diameter milling bit was used at 4000 RPM with very small feed rates to cut out the bow tie samples. Samples had triangular shapes at the two ends connected by a gage section measuring 200 microns in width. The thickness of the samples varied somewhat, however, the cross section area was correctly used for each sample to calculate the stress. Three types of samples were extracted along three directions: Growth, tool path and thickness directions (height, length and width). The overall length of the samples were about 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000934_acc.2008.4586847-Figure2.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000934_acc.2008.4586847-Figure2.1-1.png", + "caption": "Fig. 2.1: Dynamical quantities and constants for the double pendulum", + "texts": [ + " Strictly speaking, the additional spring-mass system is not a disturbance but a further uncertainty of the plant. Controller design is done for the system without the spring and the spring is taken into account by choosing suitable weights. A comparable industrial situation arises for the Hubble space telescope: Here, the solar arrays cause structural vibrations of the telescope and have to be suppressed by the controller (cf. [6]). Nonlinear pendulum dynamics. The dynamical quantities are the deflection angles \u03d5 and \u03c8 and the deflection z of the spring, cf. Fig. 2.1. In a complete rigorous approach, a forth variable would be needed, namely the angle \u03b1 which describes the lateral oscillations of the additional pendulum which is formed by the spring, the rod and the mass 0m . These oscillations are visible in the experiment but the effect on the double pendulum is very small if a motion of the double pendulum near a position where the rod is nearly horizontal is considered. Then the moment caused by the side motion depends on cos\u03b1 which is 1 in a first order approximation. If the dynamics are linearized, the side motion is completely decoupled from the rest of the pendulum and needs not to be considered for control purposes. For this reason, we assume 0\u03b1 = . Let (cf. Fig. 2.1) 2 2 2 1 1 1 1 2 0 0 2 2 2 2 2 2 . S S J J m l m L m L J J m l = + + + = + Define a generalized mass matrix by (with \u03b4 \u03c8 \u03d5= \u2212 ) 2 2 2 2 2 1 0 0 0 0 0 cos 0 ( ) cos cos 0 cos J m l L M q m l L J m L m L m \u03b4 \u03b4 \u03d5 \u03d5 = \u2212 \u2212 \u239b \u239e \u239c \u239f \u239c \u239f \u239d \u23a0 . Then the kinetic energy and the potential energy are (with ( )Tq z\u03c8 \u03d5= as the vector of the generalized coordinates) 2 1 1 2 2 2 1 2 1 2 ( , ) ( ) ( ) ( ) cos cos . TK q q q M q q P q m l m L g m l g cz\u03d5 \u03c8 = = \u2212 + \u2212 + The friction torque about 1D is fM \u03d5 , and the friction torque about 2D is fM \u03c8 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.48-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.48-1.png", + "caption": "Fig. 11.48 Gyro-moment at high speed (Note 11.18). Reproduced from Ref. [2] with the permission of Guranpuri-Shuppan", + "texts": [ + " This is different from the dynamic properties of the cornering force and the self-aligning torque because the cornering force is gradually generated with rolling distance, while a tire instantaneously deforms and generates camber thrust when a camber is applied to the tire. The dynamic cornering properties of tires at high speed are important because they relate to the vehicle response to rapid and small steering changes on a highway. These properties are different from the dynamic properties at low speed because not only the gyro moment induced by the wheels and tires, but also the vibration of the tire tread must be considered for the dynamic properties at high speed. Figure 11.48a shows that, when a slip angle is suddenly applied to a radial tire by steering a wheel at high speed, the tread area near the contact patch of the radial tire deforms laterally owing to the gyro-moment, and at the same time, the tread in the upper area deforms laterally in the opposite direction. The lateral displacement at the contact patch due to the gyro-moment delays the generation of the side force. The delay of the side force decreases with the increasing lateral spring rate of the tire and the decreasing mass moment of inertia. When a slip angle is suddenly applied to a bias tire, there is only partial tread deformation and the gyro-moment thus slightly affects the delay of the response in the bias tire. Figure 11.48b shows that, when a slip angle is suddenly applied to a radial tire from the road, or a flat-belt machine, at high speed, a force acts on the tire from the road. The force deforms the tire in the lateral direction and a gyro-moment in the counterclockwise direction around the vertical axis is then generated. Because the front part of the tire is twisted to the left while the back part of the tire is twisted to the right, the gyro-moment reduces the slip angle and thus increases the delay of the side force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002015_978-3-319-94346-6_6-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002015_978-3-319-94346-6_6-Figure8-1.png", + "caption": "Fig. 8. Schematic of improved UAV route path over multiple angle OHL.", + "texts": [ + " It is not enough to consider of using left or right side of OHL to inspect it. To improve power consumption and make it more adaptive additional movements will be done. Additional researches should be done in a field of electromagnetic influence on UAV during flight over OHL as its field could be used to prolong flight time or decrease it dependently from distance between UAV and voltage sources. As mentioned before, distance plays big role in power consumption and safety. It should be optimized. Situation as in Fig. 8 is also possible, but all previous rules must be taken into account. To fly bellow power line its sags must be known, but when fly above not always all data are possible to collect. The possible UAV routes are discussed and their parameters are defined, which allow to consider operational parameters of UAV. Within these calculations lots of additional information must be taken into account. One of most significant is regulation of overhead power lines about allowed distances and dimensions to ensure safety" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003681_978-3-030-04975-1_89-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003681_978-3-030-04975-1_89-Figure5-1.png", + "caption": "Fig. 5. Distribution of stress and deformations in the basic shape gear pump bodies, a; b) external gear pump rE, dRE, dAE, c; d) internal gear pump rI, dRI, dAI, e; f) gerotor pump rG, dRG, dAG.", + "texts": [ + " Based on that, the nature and values of the stress and strain in particular bodies needed to be determined and compared. While examining the models of the pumps, they were loaded with: \u2022 mechanical loads, namely torque T = 8, 16, 24, 32 Nm and screw clamping force Q = 35 Nm \u2022 hydraulic loads, namely working pressure p = 5, 10, 15, 20 MPa. 1 The license of the program number 05UWROCLAW has been made available to the FPRG [4] by Wroclawskie Centrum Sieciowo-Superkomputerowe of Wroc\u0142aw University of Technology. Figure 5 presents characteristic results of the stress and strain analysis of the basic shape gear pump bodies loaded with the total mechanical load by torque T = 16 Nm and screw clamping forces Q = 35 Nm, as well as with the hydraulic load by pressure p = 10 MPa. The bodies of all the pumps have been presented in the same way, namely in a general view (Fig. 5a, c, e) and section (Fig. 5b, d, f). In the view of the bodies, the areas of maximum stress and strain have been marked. The analysis of the state of stress shows that in the bodies of all pump types, the highest values of reduced stress r occur around the outlets. The areas have been marked in Fig. 5 with letter O both in the pump views and in their sections. The maximum stress r observed around outlets O is taken as a criterion for the assessment of the pump body effort. Trying to determine the source of the stress, it is needed to return to Fig. 3 which illustrates that high pressure p works on the node including the outlet chamber, the channel, the outlet of the pump, resulting in the load concentration in that node and in high stress r. Based on the FEM analysis, the relation between stress r and pressure p loading the pump, namely r = f(p), for the three considered pump types was determined", + " The analysis of the impact of the pumps\u2019 design on the stress, however, allows for the observation that the stress values in the internal gear pump body and in the gerotor pump body are lower than in the external gear pump body where hydraulic and mechanical loads concentrate around the outlet. When assessing the state of stress, it is noted that in all basic bodies shapes, with their loading with working pressure p = 0\u201320 MPa, the stresses did not exceed the allowable values. This stress condition can be written as: rE; rI;rG [rallow \u00bc 150MPa Analysing the deformations presented in Fig. 5, it can be noticed that the bodies of all the pumps become deformed both axially (A) and radially (R). Consequently, axial deformations dA and radial deformations dR of the body occur. Axial deformations dA and radial deformations dR are assumed as the criteria for the assessment of the state of the pump body deformations. Analysis of Fig. 5a, c, e shows that in the case of all the pump types, the highest values of axial deformations dA are visible mainly in the contact areas C between the central and front bodies as well as between the central and the back bodies. Similarly, analysis of Fig. 5b, d, e shows that the highest values of radial deformations dR are visible in all cases in the central body M around the outlets of all the examined pumps. Hence, reconsideration of Fig. 3 makes it possible to observe that the cause of the axial deformations is the activity of high pressure on the front face of the front and back body of a particular pump. Based on Fig. 3, it is also noted that a cause of radial deformations dR is the impact of high pressure onto the inner surface of the central body of a particular pump" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002790_detc2018-86262-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002790_detc2018-86262-Figure2-1.png", + "caption": "FIGURE 2. PARALLEL LEAF SPRING FLEXURE IN DEFLECTED POSITION, DESIRED MOTION IS ILLUSTRATED BY GREEN, PARASITICAL MOTION IN RED", + "texts": [ + " This is expressed by the following eigenwrenches, induced wrenches and eigenstiffness. 5 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 12/09/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use W\u0302 = \u22120.987 \u22120.16 0 0 0 0 0 0 1 0 0 0 0.16 \u22120.987 0 0 0 0 0 0 0.0379 0 0.066 0.998 0.101 0.009 0 \u22121 0 0 0 0 0.004 0 \u22120.998 0.066 (22) k = diag(1.9\u00b7103 5.5\u00b7106 3.7\u00b7105 1.3\u00b7104 9.4\u00b7102 7.9\u00b7102 ) (23) Again, the mechanism is actuated in the x-direction by a force acting trough the end-effector (illustrated by the red dot in Figure 2). The desired response is again a parallel translation. The applied wrench is as follows. Wa = { 1 0 0 0 0.0081 0 }T (24) Again, using Equation 19 and Equation 10 \u03bb is determined which is used to get the resultant infinite and finite pitch twists. Tf = { 5.1 \u00b710\u22124 0 \u22128.3 \u00b710\u22125 0 0 0 }T (25) T\u03c1 = { 7.6 \u00b710\u22127 0 \u22125.4 \u00b710\u22128 0 7.5 \u00b710\u22126 0 }T (26) The induced rotation T\u03c1 is not intended and is considered parasitical, which is illustrated by the red arrow. The induced translation Tf is only partly parallel with the applied wrench; this is illustrated by the green arrow in Figure 2. The part that is perpendicular to the applied twist is considered parasitical. Note that these values are the response due to a unit force. Tf\u2016 = { 5.1 \u00b710\u22124 0 0 0 0 0 }T (27) Tf\u22a5 = { 0 0 \u22128.3 \u00b710\u22125 0 0 0 }T (28) The parasitic behavior can also be expressed by the difference in the desired degree of freedom (the applied wrench) and the actual degree of freedom. The actual degree of freedom is the eigenwrench with the lowest eigenstiffness (the first column in Equation 17). For both wrenches the linear combination vectors \u03bb have to be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001901_icoa.2018.8370514-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001901_icoa.2018.8370514-Figure4-1.png", + "caption": "Fig. 4. Cross section of induction machine 3\u03d5", + "texts": [ + "Unlike other traditional means of generating electric power where the synchronous alternator is widely used, the squirrel cage asynchronous generator currently equips a large part of installed wind turbines in the world. Most applications using the asynchronous machine are intended for motor operation (this represents one-third of the world's electricity consumption), but this machine is quite reversible (fig. 3) and its robustness and low cost due to the absence of brush collectors and sliding contacts on rings make it perfectly suited for use in extreme wind conditions. The machine has a fixed pole pair number and must therefore operate over a very limited speed range (slippage less than 2%).Fig. 4 shows the induction machine 3\u03d5 in a cross section. The assumptions considered for the induction machine model are: \u2022 Magnetic hysteresis and magnetic saturation effects are neglected. \u2022 The stator windings are sinusoidally distributed throughout the air\u2013gap. \u2022 The stator slots cause no appreciable variation of the rotor inductances with respect to the rotor position. The model of Figure 7 explicitly shows the three phases of stator and rotor. Rotor and stator three phases are given by the following equations: a,b,cs, , = , , , , (2) = (3) = (4) = ( ) ( ) ( )( ) ( ) ( )( ) ( ) ( ) (5) The input of the induction machine is the output of the gear train; the following relation gives the input torque: = (6) Electromagnetic torque Te is given by the following equation: = ( ) ( ( )) ( ) (7) Where: Rs Stator resistances; Rr Rotor resistances; LS Stator self-inductances; LR Rotor self-inductances LM Mutual inductances between the stator and rotor; \u03b8 Rotor position; \u03a6a,b,cs Stator 3 phases; \u03a6a, b, cr Rotor 3 phases; P Number of poles; Jind Rotor inertia; Khs High shaft stiffness; Dhs High shaft bearing; The IC-port field is used because equation 5 contains sine and cosine relations, which are necessary to calculate the rotor position \u03b8m" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001766_978-981-10-7212-3_3-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001766_978-981-10-7212-3_3-Figure2-1.png", + "caption": "Fig. 2 The position mounting diagram of sensors", + "texts": [ + " Nevertheless, the propagation velocity of ultrasonic in the air is affected by temperature, humidity, atmospheric pressure and other factors. Among these factors, the temperature has the greatest influence on the velocity, and temperature compensation formula is given below: S = 331.45 \u00d7 N fr ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03b8 + 273.16 273.16 r \u2212 15 \u00f01\u00de where N is the number of counts, fr is reference frequency, \u03b8 is celsius temperature, and S is the desired distance. The environment perception system is composed of five ultrasonic sensors which are installed on the main control panel, as shown in Fig. 2. For the traditional ground mobile robot, the ultrasonic sensor is mounted around the robot. Therefore, the plane wall-following technology is installing three ultrasonic sensors in the front, left and right of aerial vehicle. In this paper, the three-dimensional wall-following algorithm is put forward, and two ultrasonic sensors are installed in the lower and upper parts of aerial vehicle, respectively, for constant height control. The barometer is applied to set the height of the traditional aerial vehicle outdoors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001226_kikaic.74.2870-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001226_kikaic.74.2870-Figure3-1.png", + "caption": "Fig. 3 Finite element discretization of cue", + "texts": [], + "surrounding_texts": [ + "2872 \u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b\u304a \u3051\u308b\u30ad \u30e5\u30fc \u306e\u885d \u7a81 \u7279\u6027 \u8a55\u4fa1\n(9)\n\u56de\u8ee2\u306b\u3088\u308b\u7403\u4e2d\u5fc3\u306e\u901f\u5ea6\u30d9\u30af\u30c8\u30eb\u03c1 \u3068\u56de\u8ee2\u306e\u89d2\u901f\n\u5ea6\u30d9\u30af\u30c8\u30eb\u03b8\u306e\u5404\u6210\u5206\u306e\u9593\u306b\u306f\u6b21\u306e\u95a2\u4fc2\u304c\u3042\u308b.\n(10)\n(11)\n(12)\n(13)\n3\u30fb2 \u885d\u6483\u70b9\u3092\u8868\u308f\u3059\u89d2\u5ea6 \u885d\u6483\u70b9\u306e\u4f4d\u7f6e\u3092\u8868\u308f\n\u3059\u89d2\u5ea6 \u03b1(t),\u03b2(t)\u304a\u3088\u3073\u305d\u306e\u89d2\u901f\u5ea6a,\u03b2 \u3068\u56de\u8ee2\u306e\u89d2\u901f \u5ea6\u30d9\u30af\u30c8\u30eb\u03b8\u306e\u5404\u6210\u5206\u306b\u306f\u6b21\u306e\u95a2\u4fc2\u5f0f\u304c\u6210\u7acb\u3059\u308b.\n(14)\n(15)\n(16)\n\u5f0f(14)\uff5e(16)\u306e \u5de6 \u8fba \u306f \u56de\u8ee2 \u306b \u3088 \u308b,\u305d \u308c \u305e \u308cx,\ny'\u304a \u3088\u3073Z\u65b9 \u5411 \u306e\u885d \u6483 \u70b9 \u306e \u901f \u5ea6(\u6b63 \u78ba \u306b \u306f \u901f \u5ea6/R)\u3092\n\u03b1\u304a \u3088\u3073 \u03b2\u3067 \u8868 \u308f \u3057,\u53f3 \u8fba \u306f \u03b8x,\u03b8y\u304a \u3088\u3073 \u03b8z\u3067\u8868 \u308f \u3057\n\u3066\u3044 \u308b(\u4ed8 \u9332 \u53c2 \u7167),\u5f0f(15)\u306f,a\u3092 \u03b8x\u3068 \u03b8z\u3067\u8868 \u308f\n\u3057\u3066 \u3044 \u308b\u306e \u3067,\u3053 \u308c \u3092\u5f0f(14)\u307e \u305f \u306f(16)\u306b \u4ee3\u5165 \u3059\n\u308b \u3053 \u3068\u306b \u3088 \u308a,\u03b2 \u3092 \u03b8x,\u306e \u304a \u3088\u3073 \u03b8z\u3067\u8868 \u308f\u3059 \u3053 \u3068\u304c \u51fa\n\u6765 \u308b.\n3\u30fb3 \u89e3 \u6cd5 \u3042 \u308b\u6642 \u523bt\u306b \u304a \u3051 \u308bfx(t),fy(t)\u304a \u3088\u3073\n\u53cc')\u304c \u4e0e \u3048 \u3089\u308c \u305f\u5834 \u5408\u3092 \u8003 \u3048 \u308b.\u307e \u305a,S\u22600\u3068 \u4eee \u5b9a\u3059\n\u308b.Frx(t)\u3068Frz(t)\u306f \u5f0f(17)\u3068 (18)\u304b \u3089 \u03b4(t)\u306b\u3088 \u308a\u8868\n\u3055\u308c \u308b \u306e\u3067,\u6c42 \u3081 \u308b\u3079 \u304d\u5909\u6570 \u306fbx(t),by(t),bz(t),\u03b8x(t),\n\u03b8y(t),\u03b8z(t),s(t),sx(t),sz(t),\u03c1(t),\u03b4(t),\u03b3(t),\u03b1(t)\u304a \u3088\n\u3073 \u03b2(t)\u306e14\u500b \u3067 \u3042 \u308b.\nS\u22600\u306e \u5834 \u5408\n(17)\n(18)\n\u3053\u3053\u3067,\u03bc \u306f\u6469\u64e6 \u4fc2\u6570 \u3092\u8868 \u3057,\u7c21 \u5358 \u306e\u305f \u3081\u306b,\u52d5 \u6469\n\u64e6 \u4fc2\u6570 \u3068\u9759\u6469\u64e6 \u4fc2\u6570 \u3092\u533a\u5225\u305b\u305a\u306b0\u5b9a \u5024 \u3068\u3057\u3066\u3044 \u308b.g \u306f\u91cd\u529b\u52a0 \u901f\u5ea6 \u3067\u3042\u308b.\n14\u500b \u306e\u5909\u6570 \u306e\u5185,\u03b1(t)\u3068\u03b2(t)\u306f\u5f0f(15)\u3068(14)(\u307e\n\u305f\u306f(16))\u304b \u3089 \u03b8x(t),\u03b8y(t),\u304a\u3088\u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b. by(t)\u306f\u65e2\u77e5(by(t)=0)\u3067 \u3042\u308b.S(t),\u03c1(t),\u03b4(t)\u304a\u3088\u3073y(t)\n\u306f\u5f0f(10)\uff5e(13)\u304b \u3089sx(t),sz(t),\u03b8x(t),\u03b8y(t)\u304a\u3088 \u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b.sx(t)\u3068sz(t)\u306f \u5f0f(7)\u3068(9) \u304b \u3089bx(t),bz(t),\u03b8x(t)\u304a\u3088\u3073 \u03b8z(t)\u3067\u8868\u308f \u3055\u308c \u308b.\u3057 \u305f\u304c \u3063\u3066,5\u500b \u306e\u672a\u77e5\u6570bx(t),bz(t),\u03b8x(t),\u03b8y(t)\u304a\u3088\u3073 \u03b8z(t) \u306f\u5f0f(1),(3)\uff5e(6)\u306e5\u500b \u306e\u65b9\u7a0b\u5f0f \u304b \u3089\u89e3 \u304f\u3053\u3068 \u304c\u51fa\u6765 \u308b.\u521d \u671f\u6761\u4ef6 \u306f\u3053\u308c \u3089\u306e5\u500b \u306e\u5909\u6570\u304a \u3088\u3073\u305d\u306e\n\u6642\u9593\u5fae \u5206=0\u3067 \u3042\u308b.\u3053 \u308c \u3089\u306e\u5fae\u5206\u65b9\u7a0b\u5f0f \u306f\u6709\u9650\u5dee\u5206 \u6cd5\n(\u4f8b\u3048\u3070,Newmark\u306e \u03b2\u6cd5(\u03b2=1/6)(5))\u306b \u3088 \u308a\u89e3 \u304f\u3053\n\u3068\u304c\u51fa\u6765\u308b.\n\u3082\u3057,\u89e3 \u3044\u305f\u7d50\u679c\u6700\u521d\u306e\u4eee\u5b9a\u304c\u6210 \u7acb\u305b\u305a,S=0\u306b\n\u306a\u3063\u305f\u5834 \u5408\u306f,\u524d \u5831G)\u3068 \u540c\u69d8 \u306b\u5f0f(17)\u3068(18)\u306f \u6210 \u7acb \u3057\u306a\u3044\u305f\u3081,Frx(t)\u3068Frz(t)\u304c\u672a\u77e5\u6570 \u3068\u306a\u308b\u304c,\u7897)\u3068\n\u7897)\u304c0\u3067 \u3042\u308b\u304b \u3089,\u5f0f(7)\u3068(9)\u306b \u3088\u308abx(t)\u3068bz(t) \u306f \u03b8x(t)\u3068\u03b8z(t)\u3067\u8868 \u308f\u3055\u308c \u308b.\u3057 \u305f\u304c \u3063\u3066\u672a \u77e5\u6570 \u306e\u6570 \u3068 \u65b9\u7a0b\u5f0f \u306e\u6570\u304c0\u81f4 \u3057\u89e3 \u304f\u3053\u3068\u304c\u51fa\u6765 \u308b.\n\u306a\u304a,\u7d0410-3\u79d2 \u7a0b\u5ea6\u306e\u30ad\u30e5\u30fc \u3068\u7403 \u306e\u63a5\u89e6\u6642\u9593\u5185\u3067\na(t)=0\u3068 \u4eee\u5b9a(\u524d \u5831(1)\u306b \u3088 \u308a\u7403\u306e \u4e2d\u592e\u3092\u649e \u3044\u305f\u5834\u5408 \u306b \u306f\u63a5 \u89e6\u6642\u9593\u5185\u3067\u306f\u6b86 \u3069\u3059\u3079 \u308a\u306e\u307f\u3067\u56de\u8ee2\u306f\u7121\u8996\u3067 \u304d\n\u308b)\u3059 \u308b\u3068,\u4e0a \u8ff0\u306e\u65b9\u6cd5 \u306f\u524d\u5831(3)\u306e \u65b9\u6cd5 \u306b\u5e30\u7740\u3059 \u308b.\n4. \u4efb\u610f\u65b9 \u5411\u8377\u91cd \u3092\u53d7\u3051\u308b\u30ad\u30e5\u30fc\u306e\u5909\u4f4d\n\u56f31\u306b \u793a\u3059 \u3088\u3046\u306b,\u30ad \u30e5\u30fc \u3092\u659c \u3081\u4e0b\u65b9\u306b\u50be \u3051\u3066,(\u30ad\n\u30e5\u30fc\u306e\u8ef8\u65b9\u5411\u306b)\u7403 \u306e\u659c\u3081\u4e0a\u90e8\u3092\u649e \u304f\u5834\u5408 \u306b\u306f,\u30ad \u30e5 \u30fc\u306b\u306fx\n,y\u304a \u3088\u3073Z\u65b9 \u5411\u306e\u8377 \u91cdfx(t),fy(t)\u304a\u3088\u3073fz(t)\u304c\n\u4f5c\u7528\u3059\u308b.\u3053 \u306e\u5834\u5408\u306e\u885d\u6483\u70b9\u306b\u304a \u3051\u308b\u30ad\u30e5\u30fc\u306e\u5909\u4f4d \u3092 Cx(t),cy(t)\u304a\u3088\u3073cz(t)\u3068\u3057,\u3053 \u308c \u3089\u3092\u52d5\u7684\u5f3e\u6027 \u6709\u9650\u8981\u7d20 \u6cd5\u306b \u3088\u308a\u6c42 \u3081\u308b.\u89e3 \u6790\u306b\u306f,Nastran(6)\u3092 \u7528\u3044\u305f.\n\u88681\u306e \u89e3\u67901(Normal)\u6b04 \u306e\u6750\u6599 \u5b9a\u6570 \u3092\u6301\u3064,\u9577 \u3055 1.48m,\u5148 \u7aef\u90e8\u76f4\u5f840.013m,\u5f8c \u7aef\u90e8\u76f4\u5f840.032m\u306e \u30ad\u30e5 \u30fc\u3092\u8003 \u3048\u308b.\u56f33(a)\u306b \u30ad\u30e5\u30fc\u306e\u30e1 \u30c3\u30b7\u30e5\u5206\u5272\u56f3 \u3092\u793a\n\u3059.\u56f33(b)\u306b \u5185\u90e8\u306e\u69cb\u9020 \u3092\u793a\u3059\u305f\u3081\u306b\u65ad\u9762\u56f3\u3092\u793a\u3059. 8\u7bc0 \u70b9\u306e6\u9762 \u4f53\u8981\u7d20\u307e\u305f\u306f6\u7bc0 \u70b9\u306e5\u9762 \u4f53\u8981\u7d20(\u5168 \u8981 \u7d20\u65706363\u500b,\u5168 \u7bc0\u70b9\u65706184\u500b)\u3092 \u7528 \u3044\u3066\u5206\u5272 \u3057\u3066 \u3044\u308b.\u306a \u304a,\u306f \u308a\u67f1\u8981\u7d20 \u3092\u7528\u3044 \u308b\u3068\u3088 \u308a\u7c21\u4fbf \u306a\u65b9\u6cd5\u304c\n\u5f97 \u3089\u308c \u308b\u3068\u601d\u308f\u308c \u308b\u304c,\u4eca \u5f8c\u306e\u8ab2\u984c \u3068\u3057\u305f\u3044.\n\u885d\u6483\u529b \u3092\u96c6 \u4e2d\u8377\u91cd \u3068\u3057\u305f\u305f\u3081\u306b\u885d \u6483\u70b9 \u306e\u5909\u4f4d\u306b\u8aa4\n\u5dee\u304c\u751f \u3058\u308b\u304a\u305d\u308c \u304c\u3042\u308b\u306e\u3067,\u3053 \u308c \u3092\u5c0f \u3055\u304f\u3059 \u308b\u305f\u3081 \u306b,\u30d8 \u30eb \u30c4\u306e\u63a5\u89e6\u7406\u8ad6(7)\u3067\u4e88\u60f3 \u3055\u308c \u308b\u63a5\u89e6\u534a\u5f84\u306b\u5408 \u308f \u305b\u3066\u885d\u6483\u70b9\u4ed8\u8fd1 \u306e\u8981\u7d20\u306e\u5927\u304d \u3055\u3092\u6c7a\u5b9a \u3057\u305f.\u3059 \u306a\u308f\u3061 \u4e88\u60f3\u6700\u5927\u63a5\u89e6 \u534a\u5f84(\u304a \u3088\u305d2mm)\u306e1/2\u500d \u306e1mm\u3068", + "\u30d3 \u30ea\u30e4\u30fc \u30c9\u306b\u304a \u3051\u308b\u30ad \u30e5\u30fc \u306e\u885d \u7a81 \u7279\u6027 \u8a55\u4fa1 2873\n\u3057\u305f.\u30ad \u30e5\u30fc\u306e\u30b0 \u30ea\u30c3\u30d7\u304a \u3088\u3073 \u30d6 \u30ea\u30c3\u30b8\u306b\u304a\u3051 \u308b\u62d8\u675f \u306e\u5f71 \u97ff\u306f\u5c0f\u3055\u3044 \u3068\u3057\u3066\u524d\u5831(1)\uff5e(3)\u306b\u5023 \u3063\u3066,\u652f \u6301\u306f \u81ea\u7531 \u652f\u6301 \u3068 \u3057\u305f.\u306a \u304a,\u30ad \u30e5\u30fc \u306b\u5bfe \u3057\u3066\u306fx\u8ef8 \u304a \u3088\u3073z\u8ef8 \u306e\u65b9 \u5411\u3092\u56f3]\u3068 \u9006\u65b9 \u5411\u3092\u6b63 \u3068\u3057\u305f.\n5. \u7403 \u3068\u30ad \u30e5\u30fc\u306e\u9023\u6210\n\u56f31\u306e \u3088\u3046\u306b\u89d2\u5ea6 \u03b1\u304a \u3088\u3073 \u03b2\u3067\u8868\u308f\u3055\u308c\u308b\u7403\u306e\u659c\u3081 \u4e0a\u90e8\u306e\u70b9 \u3092,\u30ad \u30e5\u30fc \u3092\u6c34\u5e73\u304b \u3089\u03c8(=\u4e00 \u5b9a)\u3060 \u3051\u4e0b\u65b9 \u306b\u5411\u3051,\u30ad \u30e5\u30fc \u306e\u8ef8\u65b9\u5411 \u306b\u649e \u304f\u5834\u5408\u3092\u8003\u3048\u308b.\u901a \u5e38, \u30bf\u30c3\u30d7\u306b\u30c1 \u30e7\u30fc \u30af\u3092\u3064 \u3051\u3066\u6469\u64e6\u529b \u3092\u5897\u3084 \u3057,\u885d \u6483\u70b9\u3067\n\u7403 \u3068\u30ad \u30e5\u30fc \u306e\u9593\u3067\u3059\u3079 \u308a\u304c\u751f \u3058\u306a\u3044\u3088 \u3046\u306b \u3057\u3066\u649e \u304f\u306e \u3067,\u885d \u6483\u70b9 \u306b\u304a \u3051\u308b\u7403 \u3068\u30ad \u30e5\u30fc\u306e\u5909\u4f4d\u304c\u7b49 \u3057\u3044 \u3068\u4eee \u5b9a\n\u3059 \u308b.\u3053 \u306e\u4eee\u5b9a \u306b\u3088 \u308a\u6b21\u5f0f\u304c\u6210\u7acb\u3059\u308b.\n(19)\n(20)\n(21)\n\u5f0f(19)\uff5e(21)\u306e \u5de6\u8fba \u306b\u304a \u3051\u308b\u6700\u5f8c\u306e2\u9805 \u306f\u5f0f(A7)\n\uff5e(A9)\u306b \u793a \u3055\u308c\u3066 \u3044\u308b\u3088\u3046\u306b\u7403\u306e\u56de\u8ee2 \u306b\u3088\u308b\u885d\u6483\u70b9\n\u306ex,y\u304a \u3088\u3073z\u8ef8 \u306b\u5bfe\u3059 \u308b\u901f\u5ea6 \u3092\u8868\u308f\u3059.\u3053 \u308c \u3089\u306e \u5f0f\u306e\u4e2d\u62ec\u5f27\u5185\u306e\u9805 \u306e\u548c \u306f\u7403\u4e0a \u306e\u885d\u6483 \u70b9\u306ex,y\u304a \u3088\u3073Z \u8ef8\u306b\u5bfe \u3059\u308b\u901f\u5ea6 \u3067\u3042\u308b.\u30ad \u30e5\u30fc \u306b\u5bfe \u3057\u3066\u306f\u524d\u8ff0\u306e\u3088 \u3046 \u306bx\u304a \u3088\u3073z\u65b9 \u5411 \u306b\u95a2 \u3057\u3066\u306f\u9006\u65b9\u5411\u3092\u6b63 \u3068 \u3057\u3066\u3044\u308b\u306e\n\u3067Cx\u3068 \u3089\u306e\u7b26\u53f7 \u306b\u6ce8\u610f\u304c\u5fc5\u8981 \u3067\u3042\u308b.\u3053 \u308c \u3089\u306e\u5f0f\u306f,\n\u03c6=0,\u03b1(t)=0\u3068 \u7f6e \u304f\u3068\u524d\u5831(3)\u306e,\u7403 \u306b\u30b9 \u30d4\u30f3\u3092\u4e0e\u3048\u308b \u5834\u5408\u306e\u5f0f\u306b\u5e30\u7740\u3059\u308b.\n\u6642\u523bt=t,(n=0,1,\u2026;t0=0)\u306b \u304a\u3051\u308bfx(tn),fy(tn),fz(tn),\n\u03b1(tn)\u304a\u3088\u3073 \u03b2(tn)\u3092\u65e2\u77e5 \u3068\u3059 \u308b.\u6b21 \u306e\u6642\u9593\u30b9\u30c6 \u30c3\u30d7t=tn+1\n(=tn+\u25b3t)\u306b \u304a \u3051\u308b.f(tn+1),f(tn+1)\u304a\u3088\u3073fz(tn+1)\u306e\u5024 \u3092\n\u4eee\u5b9a \u3057\u3066,3\u7ae0 \u304a\u3088\u30734\u7ae0 \u306e\u65b9\u6cd5\u306b\u3088 \u308acx,cy,\u3089bx,\nbz,\u03b8x,\u03b8\u03bd\u304a\u3088\u3073 \u03b8z\u3092\u6c42\u3081,\u5f0f(19)\uff5e(21)\u306e \u5de6\u8fba \u306e\u5024,\u3059 \u306a\u308f\u3061\u885d\u6483\u70b9 \u306b\u304a \u3051\u308b\u30ad \u30e5\u30fc \u3068\u7403\u306e\u5909\u4f4d\u306e\u6b8b\n\u5dee\u3092\u6c42\u3081\u308b.\n\u3053\u308c \u3089\u306e3\u500b \u306e\u6b8b\u5dee \u306e2\u4e57 \u548c \u3092\u8a55\u4fa1\u95a2\u6570 \u3068\u3057\u3066\u30b7 \u30f3\n\u30d7 \u30ec\u30c3\u30af\u30b9\u6cd5(4)\u3092 \u9069\u7528 \u3057,\u6700 \u521d\u306e\u4eee\u5b9a \u3092\u4fee\u6b63 \u3057\u3066\u4e0a\u8a18 \u306e\u8a08\u7b97\u3092\u8a55 \u4fa1\u95a2\u6570 \u304c\u5341\u5206\u5c0f \u3055\u304f\u306a \u308b\u307e\u3067\u7e70 \u308a\u8fd4\u3059.\nfx(tn+1)\u306a\u3069\u306e\u6700\u521d\u306e\u4eee\u5b9a \u306ftn-2,tn-1\u304a\u3088\u3073tn\u306b \u304a\u3051\u308b\u5024\n\u3092 \u30e9\u30b0 \u30e9\u30f3\u30b8\u30e5\u6cd5(8)\u306b\u3088 \u308a\u5916\u633f \u3057\u3066\u5b9a\u3081\u308b \u3053\u3068\u306b\u3088 \u308a, \u7e70 \u308a\u8fd4 \u3057\u56de\u6570 \u3092\u6e1b \u5c11\u3055\u305b \u308b.\n\u3053\u306e\u8a08\u7b97 \u3092t=t1\u304b \u3089\u59cb\u3081\u3066\u885d \u6483\u529b\u304c\u5341\u5206 \u5c0f\u3055\u304f\u306a \u308b (\u63a5\u89e6 \u7d42\u4e86)\u307e \u3067\u7e70 \u308a\u8fd4\u3059 \u3053\u3068\u306b\u3088 \u308a,\u5168 \u5909\u6570\u306e\u6642\u9593 \u7684\u5909\u5316\u3092\u6c42 \u3081\u308b \u3053\u3068\u304c\u3067 \u304d\u308b.\u3055 \u3089\u306b,\u885d \u7a81\u5f8c\u306e\u7403\u306e\n\u8ecc\u9053,\u56de \u8ee2 \u304a\u3088\u3073\u3059\u3079 \u308a\u306a \u3069\u306e\u5024 \u306f,\u885d \u7a81\u5f8c\u306e\u885d\u6483\u529b \u30920\u3068 \u7f6e\u3044\u3066\u5f0f(1)\uff5e(13)\u3092 \u89e3 \u304f\u3053\u3068\u306b\u3088 \u308a\u6c42\u3081\u308b \u3053\u3068\u304c\u3067\u304d\u308b.\n6. \u89e3\u6790\u7d50\u679c\u306e\u691c\u8a3c\n6\u30fb1 \u5b9f\u9a13\u306b\u3088\u308b\u691c\u8a3c \u7403 \u306e\u659c\u3081\u4e0a\u90e8 \u3092,\u30ad \u30e5\u30fc\n\u3092\u659c\u3081\u4e0b\u65b9\u306b\u50be\u3051\u3066(\u30ad \u30e5\u30fc \u306e\u8ef8\u65b9\u5411 \u306b)\u901f \u5ea6v\u3067 \u649e \u304d,\u53f0 \u306e\u4e0a\u65b9\u304b \u3089\u9ad8\u901f\u5ea6 \u30ab\u30e1\u30e9\u3067\u7403\u306e\u904b\u52d5 \u3092\u64ae\u5f71 \u3057,\n\u305d\u306e\u7d50\u679c\u3092\u672c\u5831\u3067\u958b\u767a \u3057\u305f\u7c21\u4fbf\u306a\u89e3\u6790\u624b\u6cd5(\u4ee5 \u5f8c,\u672c \u624b\u6cd5 \u3068\u547c\u3076)\u306b \u3088\u308b\u89e3\u6790\u7d50\u679c \u3068\u6bd4\u8f03 \u3057\u305f.\n\u9ad8\u901f\u5ea6\u30ab\u30e1\u30e9\u306f\u30ce\u30d3\u30c6 \u30c3\u30af\u793e\u88fdphantom v4.2(2000 \u30b3\u30de/\u79d2,196608\u753b \u7d20)\u3092 \u7528 \u3044\u305f.\u307e \u305f,\u30ad \u30e5\u30fc\u306e\u50be \u304d\u89d2 \u03c8\u304a\u3088\u3073 \u885d\u6483\u70b9 \u306ey\u8ef8 \u65b9 \u5411\u306e\u4f4d\u7f6e \u3092\u6e2c\u5b9a\u3059 \u308b\u305f\u3081 \u306b,\u771f \u6a2a\u304b \u3089\u30cb\u30b3 \u30f3\u793e\u88fd \u306e\u30c7\u30b8\u30bf\u30eb\u30ab \u30e1\u30e9D50(610 \u4e07\u753b\u7d20)\u3067 \u9759\u6b62\u753b\u3092\u64ae\u5f71 \u3057\u305f.\u89e3 \u6790 \u306b\u5fc5\u8981\u306a \u03b2(0)\u3068\u03c6 \u306f\u3053\u308c \u3089\u306e\u5199\u771f\u304b \u3089\u76f4\u63a5\u6e2c\u5b9a \u3067\u304d\u308b\u304c,\u03b1(0)\u3068v\u306f \u5e7e \u4f55\u5b66\u7684\u95a2\u4fc2\u3092\u5229\u7528 \u3057\u3066\u9593\u63a5 \u7684\u306b\u6c42 \u3081\u305f.\u3053 \u3046\u3057\u3066\u6c42\u3081 \u305f\u5024\u306f \u03b1(0)=120,\u03b2(0)=20\u00b0,\u03c8=12\u00b0,v=1.5m/s\u3067 \u3042\n\u3063\u305f.\u89e3 \u6790\u306b\u306f\u88681\u306e \u89e3 \u67901(Normal)\u6b04 \u306b\u793a\u3055\u308c\u3066 \u3044\u308b\u6750\u6599\u5b9a\u6570(3)\u3092 \u7528 \u3044\u305f.\u307e \u305f \u524d\u5831(3)\u3068 \u540c\u69d8 \u306b,\n\u03bc=0.2\u3001Mr=0,Ms=0.022N\u30fbm\u3068 \u3057\u305f.\n\u56f34\u306f0.5\u79d2 \u9593\u306b\u304a \u3051\u308b\u7403 \u306e\u8ecc\u9053 \u306e\u5b9f\u9a13\u7d50\u679c \u3068\u89e3 \u6790\n\u7d50\u679c \u3092\u6bd4\u8f03 \u3057\u3066\u3044\u308b.\u305f \u3060 \u3057,z\u65b9 \u5411\u306e\u5909\u4f4d \u304cx\u65b9 \u5411 \u306e\u5909\u4f4d \u306b\u6bd4\u3079\u3066\u975e\u5e38\u306b\u5c0f\u3055\u3044\u305f\u3081,z\u65b9 \u5411\u306e\u5909\u4f4d \u3092x\n\u65b9\u5411\u306e\u5909\u4f4d\u306e200\u500d \u306b\u62e1\u5927 \u3057\u3066 \u3044\u308b.\u5b9f\u9a13\u7d50\u679c\u306e\u8aa4\u5dee", + "2874 \u30d3 \u30ea\u30e4 \u30fc \u30c9\u306b\u304a \u3051\u308b \u30ad\u30e5\u30fc \u306e\u885d \u7a81 \u7279\u6027 \u8a55\u4fa1\n\u7bc4\u56f2\u306e\u8a18\u53f7(\u3013)\u306f,\u9ad8 \u901f\u5ea6\u30ab\u30e1\u30e9\u306b\u304a\u3051\u308b1\u753b \u7d20\u306e\u9577\u3055\u3092\u8868\u308f\u3059.\n\u3053\u306e\u56f3\u304b\u3089\u82e5\u5e72\u306e\u8aa4\u5dee\u3092\u542b\u3080\u3082\u306e\u306e\u89e3\u6790\u7d50\u679c\u306f\u5b9f\n\u9a13\u7d50\u679c\u3068\u5b9a\u91cf\u7684\u306b\u3088\u304f0\u81f4 \u3057\u3066\u3044\u308b\u3053\u3068\u304c\u308f\u304b\u308b.\u307e \u305f,\u885d \u7a81\u5f8c\u3057\u3070\u3089\u304f\u306e\u9593\u7403\u304c\u30ad\u30e5\u30fc\u306e\u65b9\u5411\u3059\u306a\u308f\u3061\u649e \u3044\u305f\u65b9\u5411\u304b\u3089\u308f\u305a\u304b\u306b\u9038\u308c\u305f\u5f8c,\u6e7e \u66f2\u3057\u3066\u3044\u308b\u3053\u3068\u304c\n\u308f\u304b\u308b.\n6.2 \u885d \u7a81\u6761\u4ef6 \u3068\u8ecc\u9053\u306e\u95a2\u4fc2 \u56f35\u306f \u7a2e \u3005\u306e \u03b1(0)\n\u3068 \u03c6\u306b\u5bfe\u3059 \u308b\u7403\u306e\u8ecc\u9053 \u3092\u6bd4\u8f03 \u3057\u3066\u3044\u308b.\u3044 \u305a\u308c\u306e\u5834\u5408 \u3082F1.5m/s,\u03b2(0)=20\u00b0 \u3068\u3057\u3066\u3044 \u308b.\u6bd4 \u8f03\u306e\u305f\u3081,\u8ecc \u90536\u306b \u56f34\u306e \u89e3\u6790\u7d50\u679c \u3092,\u8ecc \u90531\u306b \u524d\u5831(3)\u306e\u89e3\u6790\u7d50\u679c \u3092\u305d\u308c \u305e\u308c\u793a\u3059.\n\u3053\u306e\u7d50\u679c \u306f,\u03b1(0)\u3068 \u03c6\u3092\u5c0f \u3055\u304f\u3059 \u308b\u3068\u8ecc\u9053\u304c\u3042\u307e \u308a\n\u6e7e\u66f2 \u3057\u306a \u304f\u306a\u308b\u3053\u3068\u3092\u793a \u3057\u3066\u304a \u308a,\u5b9f \u969b \u306e\u7d4c\u9a13 \u30680\u81f4 \u3059\u308b\u3053 \u3068\u304c\u308f\u304b \u308b.\u524d \u5831(3)(\u8ecc \u90531)\u3067 \u306f,\u30ad \u30e5\u30fc\u3092\n\u6c34 \u5e73\u306b \u3057\u3066\u7403\u306e\u771f\u6a2a(\u03b1(0)=0,\u03c6=0)\u3092 \u649e\u3044\u305f\u5834 \u5408\u306e\u89e3 \u6790 \u3068\u5b9f\u9a13 \u3092\u884c\u3044,\u8ecc \u9053\u306fx\u65b9 \u5411\u304b \u3089\u9038\u308c\u308b\u304c,\u307b \u307c\u76f4 \u7dda \u306b\u306a\u308b\u3053\u3068\u304a \u3088\u3073,\u89e3 \u6790\u7d50\u679c \u3068\u5b9f\u9a13\u7d50\u679c\u306f\u5b9a\u91cf\u7684\u306b \u307b\u307c\u4e00\u81f4\u3059 \u308b\u3053\u3068\u3092\u793a \u3057\u305f.\u56f35\u306e \u89e3\u6790\u7d50\u679c\u306f\u524d\u5831 \u306e \u3053\u306e\u7d50\u679c \u3068\u3082\u7b26\u5408 \u3057\u3066\u3044\u308b\u3053 \u3068\u304c\u308f\u304b \u308b.\n7. \u30b7\u30e3\u30d5 \u30c8\u306e\u6750\u8cea\u304c\u885d \u7a81\u7279\u6027 \u306b\u4e0e\u3048 \u308b\u5f71\u97ff\n\u88681\u306b \u793a\u30595\u7a2e \u985e\u306e\u30b7\u30e3\u30d5 \u30c8\u306e\u6750\u8cea\u3092\u7528\u3044,\n\u03b1(0)=15\u00b0,\u03c6=15\u00b0 \u3068\u3057\u3066\u89e3\u6790 \u3092\u884c \u3063\u305f.\u4ed6 \u306e\u89e3 \u6790\u6761\u4ef6\n\u306f6\u30fb1\u7bc0 \u3068\u540c \u3058\u3067\u3042 \u308b.\u88681\u306e5\u7a2e \u985e\u306e\u6750 \u8cea\u306f,\u524d\n\u5831(3)\u3068 \u540c\u69d8 \u306b,\u6b21 \u306e\u3088 \u3046\u306b\u9078\u629e \u3057\u305f.\u89e3 \u67901\u3092 \u57fa\u6e96 \u3068\n\u3057\u3066\u89e3\u67902(Soft)\u304a \u3088\u3073\u89e3\u67903(1\u3067 \u306f\u305d\u308c\u305e\u308c \u30b7\u30e3\u30d5 \u30c8\u306e\u7e26\u5f3e\u6027\u4fc2\u6570 \u3068\u5bc6\u5ea6 \u30921/2\u304a \u3088\u30732\u500d \u3068\u3057\u3066 \u3044 \u308b.\u89e3 \u67901\uff5e3\u3067 \u306f\u5fdc\u529b\u6ce2\u306e\u4f1d\u64ad\u901f\u5ea6 \u306f\u540c0\u3068 \u306a\u308b.\n\u89e3\u67904(Light)\u304a \u3088\u3073\u89e3\u67905(Heavy)\u306f \u305d\u308c\u305e\u308c \u30b7\u30e3 \u30d5 \u30c8\u306e\u5bc6\u5ea6 \u30921/2\u304a \u3088\u30732\u500d \u3068\u3057\u3066\u3044\u308b.\u3053 \u308c \u3089\u306e\u89e3\n\u6790\u3067\u306f\u5fdc \u529b\u6ce2 \u306e\u4f1d\u64ad\u901f\u5ea6\u304c\u8868\u306e\u6700\u5f8c \u306e\u5217\u306e \u3088 \u3046\u306b\u7570\u306a \u308b.\n\u56f36(a)\u304a \u3088\u3073(b)\u306b \u305d\u308c\u305e\u308c\u89e3 \u67901(Normal),\n2(Soft),3(1\u304a \u3088\u3073\u89e3\u67901(Normal),4(Light), 5(Heavy)\u306b \u5bfe\u3059 \u308b\u885d \u6483\u529b(fx(t),fY(t)\u304a \u3088\u3073fz(t))\nball tlajectohes" + ] + }, + { + "image_filename": "designv11_92_0003335_phm-chongqing.2018.00209-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003335_phm-chongqing.2018.00209-Figure4-1.png", + "caption": "Fig. 4. The fault bearing and the fault gear", + "texts": [ + " EXPERIMENT AND SIGNAL PROCESSING In order to verify the effectiveness of the proposed method, the test was carried out on the ZJS50 comprehensive design test bench. Without affecting the normal use of the gear and bearing, the gear is manually cut off one tooth to simulate the gear failure. The gear installed on the input shaft. A small groove with a width of 0.5mm and depth of 0.2mm is processed in the outer ring of the bearing to simulate the outer race failure of the bearing. The bearing is N1007 (the number of rolling element Z=16, the contact Angle = 0, the diameter of rolling element d=6.75mm, and the diameter of pitch circle d= 47.5mm), in Fig. 4. The speed of Input shaft slowing down from 700 rev/min. The data acquisition card is NI USB9234, sampling frequency is 25.6 kHz, sampling time 50 s, charge amplifier is DH - 5853, the type of the acceleration sensor for DH112, sensitivity 5.20 pC/g. The position of acceleration sensor and eddy current displacement sensor is shown in Fig. 5. The fault feature order of N1007 bearing is computed as follows: 1 1(1 cos ) 6.88 2outer r r Z d f f f D \u03b1= \u2212 = (8) Where, fr1 is the rotation frequency of the bearing inner race" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000653_iciea.2009.5138651-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000653_iciea.2009.5138651-Figure1-1.png", + "caption": "Figure 1. Integrated design of voice coil motor and flexure mechanism", + "texts": [ + " Based on reference [12], this paper presents an integrated design consisting of a voice coil motor and flexure mechanism, to remove the highly nonlinear properties within a positioning mechanism for an operation range of more than 1mm. Moreover, a model reference adaptive control is then applied to achieve high-precision position control by using a fast adaptation mechanism to eliminate the influence of the system parameter uncertainties. Simulation results will be described to demonstrate the effectiveness of the proposed MRAC strategy and the integration of the linear actuator and flexure mechanism. II. SYSTEM MODELING The experimental setup shown in Fig. 1 is the integrated design of voice coil motor and flexure mechanism. This design can be used as an ultra-precision positioning system. The voice coil motor stands as the linear actuator, and the flexure mechanism [12] designed from prismatic joints provides the positioning mechanism, taking into consideration the deformation of the flexure member. The motion is obtained from the deformation of the flexure members, controlled by the linear actuator without any sliding and rolling contacts so that the nonlinear properties of a traditional joint, such as backlash, friction, etc", + " In the face of system parameter uncertainties such as stiffness, armature resistance, equivalent mass, and so on, the MRAC controller design, based on Lyapunov stability analysis, was applied to maintain final position accuracy under the influence of system parameter variations. The simulation results verify the effectiveness of the proposed control scheme and position mechanism to achieve a precision position better than 1\u03bcm. In future, the proposed scheme will be applied to the experimental setup shown in Fig. 1, to validate performance in practical applications. ACKNOWLEDGMENT The author would like to thank the National Science Council, Republic of China (Taiwan) for financial support of this research under grant NSC 96-2221-E-230-006. REFERENCES [1] R.H.A. Hensena, M.J.G. van de Molengraft, and M. Steinbuch, \u201cFriction induced hunting limit cycles. A comparison between the LuGre and switch friction model,\u201d Automatica, vol. 39, pp. 2131-2137, 2003. [2] J. Mao, H. Tachikawa, and A. Shimokohbe, \u201cDouble-integrator control for precision positioning in the presence of friction\u201d, Precision Engineering, vol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.33-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.33-1.png", + "caption": "FIGURE 16.33 Illustration of (a) joint gap and (b) beam misalignment.", + "texts": [ + " Also, dimplers are required for the thinner sheet of multigage blanks, to ensure even stacking on the pallet. 2. Desired length of weld. This determines the gantry specifications. 3. Number of welds required per part. A typical work cell for laser-welded tailored blank manufacturing is illustrated in Fig. 16.32. 16.8.2.4 Formability of Tailor-Welded Blanks Formability is a measure of the ease of forming a given material or body without failure. The formability of welded blanks is influenced by weld parameters such as the following: 1. Butt joint gap, which is the space between the two sheets (Fig. 16.33a). 2. Beam misalignment, which is the offset of the centerline of the laser beam from the centerline of the gap (Fig. 16.33b). 3. Shielding gas type and flow rate. 4. Focal point location. 5. Welding speed. 6. Laser power. Some observations regarding the influence of process parameters on formability may be summarized as follows: 1. Formability increases with increasing welding speed, since the increased welding speed results in a reduced weld section, even though the hardness of the weld increases. 2. An increase in misalignment or joint gap reduces formability. 3. Misalignment or joint gap tends to promote high-weld concavity, and these have relatively high failure rates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003653_ica-symp.2019.8646295-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003653_ica-symp.2019.8646295-Figure1-1.png", + "caption": "Figure 1. Two-link robotic system and schematic diagram", + "texts": [ + "05\u03b41), |\u03b41| \u2264 1 z2 = z2,0(1 + 0.05\u03b42), |\u03b42| \u2264 1. (4) Let \u03b4 = [\u03b41, \u03b42]T be the normalized uncertainty vector. The model nominal parameters can be measured as z1,0 = 0.0308 kg m2, z2,0 = 0.0106 kg m2, z3 = 0.0095 kg m2, z4 = 0.20878 kg m2, z5 = 0.063 kg m2. To verify the control performance of the proposed robust controller with LMI pole placement, a testing model, two-link robotic manipulator with parametric uncertainties is presented in Section V. The robotic system in this research is shown in Fig. 1. In this paper, two 24V DC motors are used to drive the two-link robot system as shown in Fig. 1, both DC motors can produce output torque up to 3 N m. Fig. 2 shows the block diagram of nonlinear geometric control system which consists of inner/outer loop control structure. In this section, the linearization of two-link robotic system based on a multivariable nonlinear system with parametric uncertainties is focused [7]. According to two-link robotic systems Equation (1), the form of a nominal system with state dependent perturbations can be written as follows: x\u0307 = f0(x) + \u03b4f(x, \u03b4) + g0(x)u+ \u03b4g(x, \u03b4)u y = [ y1 y2 ] = [ h1(x) h2(x) ] , (5) where f0(x) = [[ 02\u00d72 I2\u00d72 ] x \u2212D\u221210 (x)P0(x) ] , g0(x) = [ 02\u00d72 D\u221210 (x) ] , \u03b4f(x, \u03b4) = [ 02\u00d71 \u2212(D\u22121(x, \u03b4)P (x, \u03b4)\u2212D\u221210 P0(x)) ] , \u03b4g(x, \u03b4) = [ 02\u00d72 (D\u22121(x, \u03b4)\u2212D\u221210 (x)) ] , yi(x) = hi(x) = xi, and u = \u03c4 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002232_tee.22777-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002232_tee.22777-Figure1-1.png", + "caption": "Fig. 1. Structure of the magnetic levitation positioning stage", + "texts": [ + " Hussein designed a PSO-based fuzzy logic controller for the magnetic levitation system to find the optimal range and shape for the membership functions and achieved better performance than other controllers [19]. This paper proposes a kind of siding mode control strategy based on switching gain regulation by applying fuzzy logic. The parameters of fuzzy logic system are optimized by PSO, which can not only compensate the external disturbance to realize good dynamic performance but also decrease the chattering of sliding mode controller. 2. Structure of Magnetic Levitation Positioning Stage The structure of magnetic levitation positioning stage is shown in Fig. 1. It mainly consists of the planar coil, a HALBACH permanent array, and other supporting components. The planar coil is stuck on a base that is fixed onto a supporting frame after insulation. The HALBACH permanent magnet array is connected to the moving plate whose cross-section is convex and bottom is in the groove of the supporting frame in order to avoid the relatively large deviation in the nonmoving direction. When a current is passed through the coils, the electromagnetic force between the HALBACH array and the planar coil can make the stage suspended and drive it to move in the horizontal direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000806_icis.2009.184-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000806_icis.2009.184-Figure4-1.png", + "caption": "Figure 4. Membership functions for rudder angle ( c)", + "texts": [ + " In practical applications, the control goals and system constraints are all of fuzzy characters, in order to unify them, fuzzy membership function is used to express their characters. These operators can be used to translate a linguistic description of control goals into a decision function. In this way, various forms of aggregation can be chosen giving greater flexibility for expressing the control goals. The universe of discourse (range) of the inputs and outputs are mapped into several fuzzy sets of desired shapes. The membership functions for the inputs are shown in Fig 2,3 and outputs are shown in Fig 4. A fuzzy system is characterized by a set of linguistic statements based on expert knowledge. The expert knowledge is usually as \u201cif-then\u201d rules, which are easily implemented by fuzzy conditional statements in fuzzy logic. Fuzzy control rules have the form of fuzzy conditional statements that relate the state variables in the antecedent and process control variables in the consequence. Rules that were developed in the work are given in Table1. The output is defuzzified to get a final value of the control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001775_3192975.3192990-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001775_3192975.3192990-Figure2-1.png", + "caption": "Figure 2. Schematic representation of the Qball-X4.", + "texts": [ + "i PWM input ui by a first-order linear transfer function as follows: ii u s kT (1) where i =1, 2, 3, 4 and k is a positive gain and \u03c9 is the motor bandwidth. k and \u03c9 are theoretically the same for the four motors but this may not be the case in practice. It should be noted that ui=0 corresponds to zero thrust and ui=0.05ms corresponds to the maximal thrust that can be generated by the No.i motor. The block diagram of the UAV system is illustrated in Fig. 3. The model of quad-rotor UAV is shown in figure 2. Four rotors which generated propeller forces F1, F2, F3 and F4 are used to simplify the displacement and to increase the lift force. A possible combination of controlled outputs can be the Cartesian position (x, y, z) and the yaw angle \ud835\udf13 in order to track the Permission to make digital or hard copies of all or part of this work for personal or classroom use is granted without fee provided that copies are not made or distributed for profit or commercial advantage and that copies bear this notice and the full citation on the first page", + "3192990 desired position trajectory, more to an arbitrary heading and stabilize the pitch and the roll angles (\ud835\udf19, \u03b8), which introduces stable zero dynamics into the system. The motors and propellers of a quad rotor are configured in such a way that the back and front (1 and 2) motors spin clockwise (thus inducing two counterclockwise torques on the body) and the left and right (3 and 4) spin counterclockwise (thus inducing two clockwise torques on the body). Each motor is located at a distance L from the center of mass o (0.2 m) and when spinning, a motor produces a torque \ud835\udf0f\ud835\udc56 which is in the opposite direction of motion of the motor as shown in Fig. 2. The origin of the body-fixed frame is the system's center of mass o with the x-axis pointing from back to front and the yaxis pointing from right to left. The thrust Ti generated by the i th propeller is always pointing upward in the z-direction in parallel to the motor's rotation axis. The thrusts Ti and the torques \u03c4i result in a lift in the z-direction (body-fixed frame) and torques about the x-, y- and z-axis. The relations between the lift/torques and the thrusts are 4311 43 21 4321 u TTLu TTLu TTTTuz (2) The torque \u03c4i produced by the ith motor is directly related to the thrust Ti via the relation of ii TK with K as a constant" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003050_jae-171190-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003050_jae-171190-Figure1-1.png", + "caption": "Fig. 1. The electromagnetic pump developed.", + "texts": [ + " As a result, it is possible to reduce the rotation torque of the magnet and use a small motor. Therefore, reduction in size of the entire apparatus is expected. It is also expected to reduce power consumption. In this research, we develop linear actuator with features suitable for artificial hearts with such advantages. In addition, as a preliminary step to the application to the artificial heart, we applied it to the electromagnetic pump and decided to evaluate the basic structure and characteristics of the experiment. Figure\u00a01 shows the pulsating electromagnetic pumps developed using permanent magnet linear actuator. Figure\u00a02 shows 3D-CAD model of the electromagnetic pump. The electromagnetic pump newly developed in this research is mainly composed of permanent magnet, tank, driving part. The drive uses a DC motor, and the reduction ratio between the motor and the permanent magnet is 1:180. We use 4 poles, 6 poles, 8 poles for permanent magnets. Each magnet reciprocates 90 degrees for 4 poles, 60 degrees for 6 poles and 45 degrees for 8 poles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000848_vss.2008.4570702-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000848_vss.2008.4570702-Figure2-1.png", + "caption": "Fig. 2. Typical trajectories with g\u2217 = \u22121. Left: s(0)s\u0307(0) > 0.", + "texts": [ + " This means, in light of (8) and (9), that it cannot exist t\u2217 > 0 enabling the sign inversion of \u03c1(t). Thus the system will reach the origin of the s \u2212 s\u0307 plane in finite time [1]. g\u2217 = \u22121. It must be proven that if g\u2217 = \u22121 there is a finite instant t\u2217 at which |s| exceed the threshold value \u0394. The closed loop dynamics is described by the differential inclusion s\u0308 \u2208 [\u2212F, F ] \u2212 [G1, G2]u(t) (13) with, by construction, \u03c1(t) = 1 when t \u2208 [0, t\u2217]. The transient behaviour is qualitatively different depending on the sign of the product s(0)s\u0307(0). If s(0)s\u0307(0) > 0 (see Fig. 2, left plot) then sign u(t) = \u2212sign(s(0)) = \u2212sign(s\u0307(0)) until a singular point is achieved. This means that s\u0308 will always have the same sign of s(0) and s\u0307(0). This implies that |s| and |s\u0307| are strictly increasing functions, which implies, in turns, that in finite time t\u2217, |s| will overcome the value \u0394. If s(0)s\u0307(0) < 0 the transient trajectories can be more variegated (see Fig. 2-right). Now we have that sign u(t) = \u2212sign(s(0)) = sign(s\u0307(0)) in the initial part of the trajectory. If a new singular point is achieved without hitting the line s(t) = \u03b2s(0) (curve (a)) then |s| will continue to grow afterwards, and, therefore, the threshold |s| = \u0394 will be exceeded in finite time. If the line s(t) = \u03b2s(0) is hit before to achieve the new singular point (curve (b)) the control will reverse its sign after hitting the line. Since, at this time, the sign of s\u0307(t) has not yet changed (no new singular point has been still achieved) this means that the sign of u will never change afterwards, which implies that in finite time t\u2217, |s| will overcome the value \u0394" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.15-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.15-1.png", + "caption": "Fig. 11.15 Deformation and force of a tire in the contact area. Reproduced from Ref. [2] with the permission of Guranpuri-Shuppan", + "texts": [ + " Considering the left\u2013right symmetry of the lateral displacement y, the relation dy/dx = 0 is satisfied. Using the condition that the integration of the reaction of the carcass spring must be equal to the side force Fy, y is given by4 4Note 11.2. y \u00bc kFy 2ks e kx cos kx\u00fe sin kx\u00f0 \u00de; \u00f011:39\u00de where k is expressed by k \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ks=\u00f04EIz\u00de4 p : \u00f011:40\u00de Applying the Taylor series expansion to Eq. (11.39) with respect to kx, the lateral displacement y can be expressed by a parabolic function5: y \u00bc kFy 2ks 1 k2x2 : \u00f011:41\u00de Figure 11.15 shows the deformations of the tread base (belt) and the tread surface near the contact patch when a tire rolls with slip angle a. The point on the road surface moves on the straight line ABD while the point on the tread base moves on the line AC. The point of the tread surface makes contact with point A of the road and moves straightly to sliding point B, subsequently sliding to point C. Section AB is the adhesion region, and section BC is the sliding region. Suppose that the deformation of the beam in the contact area is expressed in the coordinate system (x1, yb) where the position of the belt at the leading edge is the origin", + " When the tread ring is twisted by the self-aligning torque Mz, the slip angle is decreased by the twisting angle. The effective slip angle a is thus expressed by a \u00bc a0 Mz=Rmz; \u00f011:68\u00de where a0 is the slip angle at the wheel. 11.2.3 Kinetic Friction Coefficient of Rubber Changing with Sliding Speed The kinetic friction coefficient of rubber ld decreases with sliding speed V 0j j: ld \u00bc ld0 aV V 0j j; \u00f011:69\u00de 9Note 11.7. where ld0 is the static friction coefficient of rubber and aV is a constant that expresses the speed dependency of the friction coefficient. Referring to Fig. 11.15 where only the slip angle is applied to a tire, while the point on the road surface moves from point B to point D with speed V, the point on tread moves from point D to point C. Sliding occurs in the sliding region from lh to l, and the average sliding speed V 0j j in the sliding region is then given by V 0j j \u00bc lV sin a=\u00f0l lh\u00de; \u00f011:70\u00de where V is the tire speed. The substitution of Eq. (11.70) into Eq. (11.69) yields ld \u00bc ld0 aV lV sin a=\u00f0l lh\u00de: \u00f011:71\u00de The shape of the contact pressure distribution of a tire is different from the parabola used in the Fiala model, and Sakai [2] thus proposed expressing the distribution using a generalized parabolic function: qz \u00bc n\u00fe 1 n 2nFz ln\u00fe 1b l 2 n x1 l 2 n ; \u00f011:72\u00de where b is the contact width, l is the contact length, Fz is the load, and x1 is measured from the leading edge of the contact patch. n is a parameter that controls the pressure distribution; n = 4 is appropriate for a passenger-car radial tire. When a tire rolls with slip angle a, the relative lateral displacement Dy is defined in a coordinate system having its origin located at the leading edge as shown in Fig. 11.15. Using Eq. (11.67), the relative lateral displacement Dy at distance x1 is given by Dy \u00bc x1 tan a yb \u00fe yc\u00f0 \u00de \u00bc x1 tan a d Cy \u00fe l2 3r2Ky Fy x1 l 1 x1 l : \u00f011:73\u00de The sliding point lh can be obtained from the equilibrium between the shear force and the maximum frictional force. Using Eqs. (11.72) and (11.73), lh is determined as n\u00fe 1 n 2nFzls ln\u00fe 1b l 2 n x1 l 2 n \u00bc Cy x1 tan a d Cy \u00fe l2 3r2Ky Fy x1 l 1 x1 l : \u00f011:74\u00de Equation (11.74) can be easily solved using an iterative procedure. In the case that lh < 0, lh = 0 is used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003929_0954406219843951-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003929_0954406219843951-Figure3-1.png", + "caption": "Figure 3. Pinion shaper cutter and basic rack.", + "texts": [ + " Obviously, the standard basic rack is a special case to the proposed design by setting r0,t, \u00bc r0,t, . 15 Besides that the tool must be conjugate to this basic rack, the tool tooth number zc and possible addendum correction xcmn must be known. The tooth profile is found by rolling the basic rack over the tool\u2019s pitch circle. The tool has involute profile in, and only in, the transverse plane. In this plane, the basic rack coordinates are t \u00bc n cos t \u00bc n \u00f04\u00de It is suitable to describe the tooth of the shaper cutter. In Figure 3, one tool tooth share the contact point C with the basic rack. The contact normal in every contact must be directed through the pitch point P. This gives the coordinates of the tool tooth profile. rc n,c, c \u00bc c n, c\u00f0 \u00de c n, c\u00f0 \u00de c 0 B@ 1 CA \u00bc R0,t sin \u00fe n\u00fexc sin \u2019 cos \u00fe \u2019\u00f0 \u00de R0,t cos \u00fe n\u00fexc sin \u2019 sin \u00fe \u2019\u00f0 \u00de c 0 B@ 1 CA \u00f05\u00de where c\u00f0 n,c, 0\u00de divides the tooth in two equal symmetric parts. Here cot \u2019 \u00bc d n,c d n,c cos \u00f06\u00de and n,c, c \u00bc n,c cos \u00fe n,c \u00fe xc cot \u2019 c tan R0,t \u00f07\u00de The shaper cutter removes material by cutting planes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure7-1.png", + "caption": "Fig. 7. Third metamorphic limb with phase e2", + "texts": [ + " Therefore, phase e2 of metamorphic joint conform motion type of submanifold T2(w3) \u00b7 S(p3) and phase e1 of metamorphic joint coincide motion type of submanifold T2(v3) \u00b7 S(q3) by combining the above constraint conditions. Therefore, transformation of motion type from submanifold T2(w3) \u00b7S(p3) to submanifold T2(v3) \u00b7S(q3) is correspond with transformation from phase e2 to phase e1 of metamorphic joint. 2) Type synthesis of third metamorphic serial limb: Submanifold T2(w3) \u00b7S(p3) limb has been generated by one of R(p31,u) \u00b7R(p31,v) and R(p32,w3) \u00b7R(p33,w3) \u00b7R(p34,w3) with respect to local coordinate system o3 \u2212 u3v3w3. The metamorphic limb is synthesized with the phase e2 in Fig. 7. Link a is fixed to base, and rotational joint of axis 1 is locked. Axis 2 and axis 3 are perpendicular each other which is equivalent to Hooke joint with u and v rotate axis direction. Axis 4 and axis 5 are parallel to axis 6 with w3 direction. Then metamorphic limb takes on mobility five which can realize motion type of submanifold T2(w3) \u00b7S(p3). Keeping axis 1 is locked and rotating link c so that axis 2 and axis 3 with u and w rotate axis direction in Fig. 8. Axis 4 and axis 5 are parallel to axis 6 with v3 direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003373_imece2018-86310-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003373_imece2018-86310-Figure1-1.png", + "caption": "Fig. 1: The schematic of the Gantry-type 3D printer (left) and the image of the actual printer (right).", + "texts": [ + " Image analysis of landmarks on the samples is used to obtain the strain. The 3D printed metallic sample was obtained from a solid rectangular walled structure welded by a 3D printer. The torch of a MIG welder was attached to the carriage of a gantry printer to created layers of weldment on top of each other. The movement of the torch was controlled by a MACH3 program in x, y, and z direction. The arc welding process was performed without shield gas. To illustrate the process, a picture of the 3D printer and its schematic are shown in Fig. 1. The movement of the torch is depicted in Fig. 2 along with the directions including tool path, growth and thickness. 2 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 01/21/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use Once solidified, the sample was removed from the base plate and machined with a CNC to produce bow-tie samples shown in Fig. 3. A 1mm diameter milling bit was used at 4000 RPM with very small feed rates to cut out the bow tie samples" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003738_978-3-030-04975-1_7-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003738_978-3-030-04975-1_7-Figure2-1.png", + "caption": "Fig. 2. Rolling pairs of the cycloidal gear transmission", + "texts": [ + " In one of them, the epicycloidal curve makes the contour of the external gear, whereas in the other one, the hypocycloidal curve is the outline of the internal gear. The transmission with the hypocycloidal gearing is shown in Fig. 1 [8]. The cycloidal gears and the rollers form the internal gearing. The cycloidal transmissions are mainly used in drive systems as reducers. A cycloidal gear transmission is a rolling transmission in which all geometrically connected elements are moved by rolling motion (Fig. 2). This results in a maximum reduction of losses caused by friction. Three rolling pairs can be distinguished: \u2022 planetary gear (1) with rollers (2) collaborates with fixed body (6) with the hypocycloidal profile; rollers (2) rotate around their own axis at speed xr; \u2022 planetary gear (1) collaborate with the central bearing (of the eccentric) (3); the gear is fixed eccentrically on the input shaft, moving at angular speed xh; \u2022 planetary gear (1) collaborates with sleeve (5), together with pin (4) moving at speed xt in the hole of the planetary gear", + " The power loss of the rolling friction occurring between the roller and the centre of the planetary gear can be defined as: NTk r \u00bc Fi fk r xk \u00fexr\u00f0 \u00de \u00f02\u00de where: fk-r \u2013 coefficient of the rolling friction between the roller and the planetary gear centre. The power loss of the rolling friction occurring between the roller and the hypocycloidal outline of the body is defined as: NTr o\u00f0hipo\u00de \u00bc Fi fr o xr \u00f03\u00de where: fr-o \u2013 coefficient of the rolling friction between the roller and the outline of the hypocycloidal body. The angular speed of the i-th roller xr fixed in the planetary gear can be determined using Fig. 2. On its basis, it is possible to define the relationship of the angular velocity of gear xk with the angular velocity of roller xr: xr \u00bc ui 2 g\u00f0 \u00de 2 g xk \u00f04\u00de where: ui \u2013 the distance from the contact point of the roller and the toothing to rolling point C of the planetary gear; ui = AiC. The relation of the angular velocity of the planetary gear xk and the speed of the input shaft xh is: xk \u00bc xh e rw1 \u00bc xk u \u00f05\u00de where: e \u2013 eccentricity, rw1 \u2013 rolling radius of the planetary gear, u \u2013 gear ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001699_1077546318767559-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001699_1077546318767559-Figure3-1.png", + "caption": "Figure 3. Dynamic model of a helical gear pair with mesh excitation.", + "texts": [ + " KiLand KiR are the bearing stiffness matrix of the left bearing and the right bearing of shaft i, while CiL and CiR are the support damping matrix of the left bearing and the right bearing, respectively. The geared rotor system shown in Figure 2 can be divided into different elements to be modeled and analyzed by the finite element method. The corresponding gear nodes are connected by gear elements, and every shaft is supported by two bearing elements. The bullgear and other additional parts of the counterweight plate fixed on the shaft are simplified into a lumped mass element. The dynamic model of the helical gear is shown in Figure 3, where i and j are driving and driven gears, respectively, while !i and !j also can be expressed by _ xi and _ xj. Meshing dynamic excitation of gear pairs mainly includes the time-varying mesh stiffness, transmission error, and backlash. 1. Time-varying mesh stiffness In the transmission process, the mesh stiffness has periodicity, and can be expressed in the form of a Fourier series: k\u00f0t\u00de \u00bc k0 \u00fe X1 i\u00bc1 ai cos\u00f0i! t\u00de \u00fe bi sin\u00f0i! t\u00de\u00bd \u00bc k0 \u00fe X1 i\u00bc1 ki cos\u00f0i! t\u00fe \u2019i\u00de \u00f01\u00de where k0 is the average mesh stiffness, ai, bi, and ki are Fourier series parameters, \u2019j is the phase angle, and ", + "5 Tooth width (mm) 45 Pressure angle (deg) 20 Helical angle (deg) 12 Teeth number 80 193 41 33 25 Base radius (mm) 191.6 462.3 98.2 79.1 59.9 Mass(kg) 13 50 3.8 2.5 1.7 Polar moment of inertia (kg m2) 0.07 1.6 0.007 0.004 0.003 Diameter moment of inertia (kg m2) 0.04 0.8 0.005 0.003 0.002 Rated torque (N m) 153 \u2013 24 20 18 the nonlinear meshing characteristic of gears, meshing force will show complicated nonlinear characteristics, while the unbalance of rotor system will influence these characteristics. Based on the geared rotor system shown in Figure 3, the torque of the three output shafts under rated torque is 24N m, 20N m, and 18N m, respectively. The meshing forces between output shaft 1 gear and intermediate gear, output shaft 2 gear and intermediate gear, and output shaft 3 and intermediate gear are defined as F23, F24, and F25, respectively. In order to find the effect of unbalance, a series of big unbalance values are set in the simulation, as shown in Table 3. Based on the established dynamic analysis model, the meshing forces are compared through numerical calculation considering gear nonlinear factors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.29-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.29-1.png", + "caption": "FIGURE 16.29 Cadillac center pillar inner. (From Uddin, N., Photonics Spectra, 1993.)", + "texts": [ + " This section focuses on laser-welded tailored blanks. Tailored blanks have replaced a number of units that are traditionally produced by first forming single components, which are then subsequently joined by welding, especially for the automotive bodyin-white. One typical application of tailored blanks in this area is the use of a thicker or higher strength material to replace the reinforcement that would normally be required for strength or support. An example is the center pillar inner of a car body (Fig. 16.29) where a thicker material is used in the upper portion where a separate reinforcement would normally be needed and joined to the thinner material that is subjected to lower stresses, before forming. Laser-welded blank Final formed part Original-design 16.8.2.1 Advantages of Tailored Blank Welding The advantages of tailored blanks over conventionally produced body-in-white components include the following: 1. Weight reduction, achieved by welding a thinner gage to a thicker gage material that provides reinforcement only where needed and also by elimination of reinforcement panels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001893_(asce)em.1943-7889.0001473-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001893_(asce)em.1943-7889.0001473-Figure2-1.png", + "caption": "Fig. 2. Two coordinate systems. (Reprinted from Padron et al. 2011, with permission.)", + "texts": [ + " In addition, the following boundary conditions at the orifice where the jet exits (Wallwork et al. 2002; Padron et al. 2011) need to be satisfied: X \u00bc Y \u00bc Z \u00bc \u2202Y=\u2202s \u00bc \u2202Z=\u2202s \u00bc 0; \u2202X=\u2202s \u00bc 1; u \u00bc U; R \u00bc ro at s \u00bc 0 \u00f01f\u00de where U = centerline velocity of jet at exit section; and ro = radius of orifice at jet exit section. As in the experimental observation and configuration for FS (Padron et al. 2011, 2013), here a rotating jet with a curved centerline is considered, and, as in the work due to Wallwork et al. (2002) as well as Padron et al. (2011), first the rotating system (Fig. 2) is set up with two coordinate systems, where one is fixed \u00f0x; y; z\u00de and the other \u00f0X;Y;Z\u00de is attached to the rotating spinneret, and both systems have a common origin at the center of the spinneret. In Fig. 2 only a cylindrical portion of the spinneret whose half distance from the origin of the coordinate systems to the orifice at the right side is designated by the length C. However, in the actual experimental case, the fiber can emerge from the orifices at both ends of the shown cylindrical portion of the spinneret, but in the present analysis, only that fiber jet is considered that is Fig. 1. Forcespinning schematic. (Reprinted from Padron et al. 2011, with permission.) \u00a9 ASCE 04018069-2 J. Eng. Mech. J. Eng. Mech., 2018, 144(8): 04018069 D ow nl oa de d fr om a sc el ib ra ry .o rg b y U ni ve rs ity O f Sy dn ey o n 07 /1 6/ 18 . C op yr ig ht A SC E . F or p er so na l u se o nl y; a ll ri gh ts r es er ve d. generated and emerges from the orifice to the right side, as shown in Fig. 2. In this figure the magnitude of the angular velocity of the rotating spinneret is shown in a more general case as a function of time treated in Padron et al. (2011), but in the present work such angular velocity will be treated as a given constant. Also in Fig. 2, C represents the half-length of the spinneret. In the present study the analysis and computation will be based on a local orthogonal curvilinear coordinate system, where such a curvilinear coordinate system can describe the location of any point p on the fiber. The center of this local curvilinear coordinate system is considered to be a point o on the jet\u2019s centerline; further, es denotes a unit vector along the jet\u2019s centerline, p and b the principal normal and binormal vectors, respectively, to the centerline and in a plane perpendicular to es, and en and e\u03d5 the unit vectors along the radial and azimuthal directions, respectively, at point p in the fiber", + " (2002) in a related problem for the effects of rotation and surface tension on jet properties, it was found that for the present FS problem, the jet speed increases and jet radius decreases with decreasing Rb, which is reasonable since a smaller Rossby number implies larger rotational forces, which can enhance the jet speed and make the jet thinner, as observed experimentally (Padron et al. 2013). It was also found that the speed of a fiber jet decreases and its radius increases as the Weber number decreases, which is due to the force resisting the thinning of the fiber. Figs. 3 and 4 present typical results on the effects of the aerodynamic drag force on jet speed, jet radius, and trajectory. Fig. 2 shows the centerline velocity versus arc length for Rb \u00bc 0.8, We \u00bc 1.5, and different values of \u03b3 (\u03b3 \u2261 \u03b31 \u00bc \u03b32). It can be seen from this figure that jet speed decreases with increasing aerodynamic drag force, which is understandable because of the stabilizing effect of the drag force. However, jet speed increases with arc length mostly for an arc length of less than some value s0, where s0 decreases with an increasing value of aerodynamic drag. Also, the rate of increase of jet speed with respect to the arc length decreases with as the effect of aerodynamic drag increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002227_s00706-018-2202-2-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002227_s00706-018-2202-2-Figure5-1.png", + "caption": "Fig. 5 Illustration of the instrumental setup used for EC\u2013 CE\u2013MS measurements: a injection unit with novel injection cell and electrolyte reservoirs, b potentiostat, c high voltage source, d fused silica capillary, e mass spectrometer, f computer", + "texts": [ + " Due to materials and modularity of the cell, it could be cleaned easily and was suitable for measurements in aqueous as well as non- aqueous solutions. Electrodes could easily be exchanged, which allowed for high flexibility regarding electrode materials. When installed in the EC\u2013CE\u2013MS device, a fully automated hydrodynamic injection of sample directly from the working electrode surface was possible by placing the tapered tip of the fused silica separation capillary onto the electrode surface. The overall experimental setup is illustrated in Fig. 5 in the experimental section. By CE\u2013MS, a fast separation and detection of neutral and particularly cationic species was possible applying a positive high voltage at the injection end of the capillary (detection end of capillary installed in grounded ESI sprayer). Figure 2a shows a CE\u2013MS measurement that was carried out at a separation voltage of 18 kV (2.4 lA) without previous oxidation using a model mixture of Fc, FcMeOH, and dMFc. The three model substances showed different behavior regarding state of charge and detectability", + " A thin-film electrode with gold working, counter, and quasireference electrode (ED-SE1-Au, Micrux Technologies, Oviedo, Spain) was used for the oxidation. All potentials given are referred to the Au quasireference electrode and all electrochemical measurements were carried out in ACN/ 10 mmol/dm3 NH4OAc/1 mol/dm3 HOAc as BGE. The electrode was installed in the novel injection cell described in detail in the results and discussion section. A lSTAT 200 potentiostat (DropSens, Llanera, Spain) controlled by Dropview 200 software was used for applying potentials. A schematic illustration of the experimental setup is depicted in Fig. 5. For evaluation of the cell performance, a solution of 1.5 mmol/dm3 Fc, 1 mmol/dm3 FcMeOH, and 40 lmol/ dm3 dMFc in BGE was used. Fast detection studies were carried out with a solution of 1 mmol/dm3 FcMeOH in BGE. For the CE protocol, 8 mm3 of sample solution were filled into the cell. The sample was hydrodynamically injected into the CE system by placing the tapered end of the separation capillary onto the working electrode surface for 2 s at a difference in height of 18 cm between the injection end of the capillary and the detection end of the capillary (hydrostatic pressure)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002886_s1068798x18100179-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002886_s1068798x18100179-Figure5-1.png", + "caption": "Fig. 5. Structure of pipelines 1 (a) and 2 (b): (1) pipeline; (2) clip; (3) crosspiece.", + "texts": [ + " In the proposed method 3, in comparison with modal analysis (method 1) and dynamic analysis with an increasing frequency (method 2), we note significant redistribution, and correction of the eigenfrequencies, amplitudes, and stresses on account of taking account of the model configuration, the attach- ment system (in contrast to the method 1), and the absence of additional external excitation (in contrast to method 2). The reliability is estimated for a segment of the pipeline\u2019s working model. The two pipeline structures considered differ only in the f lexure radius: pipeline 1 is shown in Fig. 5a, and pipeline 2 in Fig. 5b. All the parameters remain the same as for the test model in method 3 (Table 1). In the analysis, displacing forces are applied to the pipeline supports (Fig. 5b). The displacement vectors are collinear with the Z axis. In the analysis, we obtain the frequency responses for both structures. In Fig. 6, we show the corresponding spectrograms (shock responses). Assessment of the results shows that many frequencies for pipeline 2 are shifted upward by 20\u201335 Hz in comparison with pipeline 1. Note that different vibrational responses are observed for all the structures under the action of the same force with change in only one pipeline section. We now assess the results of shock loading with change in one geometric parameter of a section of the prefabricated structure. In Fig. 7, we plot the stress against the time. For pipeline 1 (Fig. 5a) and pipeline 2 (Fig. 5b), the stress pattern is similar, but there are differences not only in the frequency characteristics but also in the amplitude and periodicity of the stresses in the structure. That will undoubtedly affect its working life. The peak stress in pipeline 2 is 15% higher. Any version of reliability theory may be used to calculate the number of working cycles to failure. The result will not predict the operation of a real structure [5]. Rather, it may be used to determine the increase in reliability of the modified structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000213_s11665-008-9314-5-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000213_s11665-008-9314-5-Figure4-1.png", + "caption": "Fig. 4 Geometry and dimensions of the U-bolts. Dimensions in mm", + "texts": [ + " Once the steels were selected and produced, a previous analysis was carried out using tensile tests to establish the amount of plastic deformation during the drawing process (Ref 4). Therefore, bars from the proposed steels were obtained by drawing process to promote a general increase of the mechanical resistance by strain hardening effect. To increase this strength even further, to this previous deformation an additional local deformationwas carried out during the cold forming of the thread regions by a special rolling process. The bolts were manufactured in a U format with the preselected threads (M129 1.25) on it (Ref 5). Figure 4 presents the geometry and dimensions of the U-bolts. Thematerial chemical analysis of the selected steels was carried out using an optical spectrometry by arc Mod 3460. For microstructural evaluation, samples were removed from the thread (root and crest) and A and B regions, on the longitudinal and transversal directions as indicated in Fig. 4. In all cases, the analyses were carried out on the surface and in the center of the bar. The removed samples were grounded, polished, and etched with 2% Nital. Finally, they were observed in an optical microscope in conjunction with a Buehler Omnimet Enterprise image analyzer for grain size (GS) evaluation on the transversal direction. From the microstructural evaluation, it was possible to verify the microstructure composition and the average GS following the ASTM E112-96 standard (Ref 6), as well as the decarburizing and discontinuities present in the steels" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003299_6.2019-1258-Figure16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003299_6.2019-1258-Figure16-1.png", + "caption": "Fig. 16 Test frame", + "texts": [], + "surrounding_texts": [ + "Based on the promising results of initial damped mechanism tests, a second prototype mechanism was built to conduct further damped deployment tests. This mechanism included a bracket to maintain alignment between axles of the hub and the roller. The bracket is labeled in Figure 1. It includes a slot so the roller can maintain close contact with the boom even as the radius shrinks during deployment. This prototype also included a gear to with the damper\u2019s pinion and drive the damper faster for increased damping. This prototype can be seen in Figure 12 D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58 0 20 40 60 80 100 120 deployment distance, cm 0 50 100 150 200 250 300 350 400 450 sp ee d, c m /s undamped damped Fig. 11 Deployment mechanism speed Fig. 12 Second prototype with bracket and gear D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58 The second prototype was used to test deployment with all three dampers. The boom deployed successfully using all three dampers with no blossoming. Total deployment times were 2.1, 7.8, and 25.7 s for the light, medium, and stiff dampers. Deployment speeds can be seen in Figure 13. These tests demonstrate successful accomplishment of all of the mechanism\u2019s primary goals even at very slow deployment speeds. 0 20 40 60 80 100 120 deployment distance, cm 0 10 20 30 40 50 60 70 sp ee d, c m /s light medium stiff Fig. 13 Boom deployment speed with three different dampers VII. Flight Testing The MUSE mechanism was also tested in freefall on a parabolic research flight flown by Zero-G. The flight trajectory, shown in Figure 14, provides approximately 17 s of freefall to create a space-representative environment [14]. Fig. 14 Parabolic flight trajectory The boom was suspended from a test frame and deployed shortly after the onset of zero-g. The test frame can be seen in Figures 15 and 16. The undamped mechanism successfully controlled the boom\u2019s deployment path, as seen in Figure 17. This deployment lasted 2.73 s, and the mechanism reached a maximum speed of just over 1 m/s. Deployment speed is shown in Figure 19, along with aircraft acceleration in the vertical direction. D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58 Fig. 15 Test frame in flight The mechanism successfully prevented blossoming during handling and transit, as well as through most of deployment. However, there is a small amount of blossoming visible at the end of deployment. This can be seen in Figure 18. For future tests we will attempt to eliminate blossoming altogether by fixing the boom more securely to the hub and by using stronger springs between the hub and the roller. Initial handling tests indicate that doubling the spring force makes the booms much more resistant to blossoming. VIII. Conclusion This paper describes the ongoing work on the MUSE deployment mechanism developed at NCSU. This deployment device is a minimal unpowered strain-energy deployment mechanism suitable for monostable booms. Testing shows that this mechanism successfully achieves all of its primary objectives. 1) Ensure reliable jam-free operation 2) Control the deployment path and eliminate blossoming 3) Control deployment speed We demonstrated some unique characteristics of this mechanism that make it suitable for space applications. A single sprung roller can create enough inter-layer friction to prevent blossoming in a coiled boom, even during a slow deployment. The system has been shown to be tunable and is easily adjustable for different length booms compared to other deployment mechanisms. Deployment speed can be tailored by using dampers of varying stiffness. The results lay the groundwork for a lightweight minimal deployment device that can be adapted to different booms types and configurations to provide controlled deployment. IX. Future work We will continue our flight test campaign with an additional flight in 2019. During this flight we will test a damped mechanism as well as booms stored at various temperatures. D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58 173 cm Fig. 17 Undamped deployment in freefall Fig. 18 Blossoming beginning near full deployment D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58 0 0.5 1 1.5 2 2.5 3 time, s 0 0.2 0.4 0.6 0.8 1 1.2 sp ee d, m /s deployment speed frame acceleration X. Funding This research is supported by NASA Flight Opportunity Grant 80NSSC18K1282 along with the Space Act Agreement SAA1-23398_Annex_3 between NCSU and NASA Langley. XI. Acknowledgments The authors would like to thank Vince Chicarelli, Steve Cameron, Gary Lofton, Bryson Jones, and Ryan Dykes for the excellent service and machining expertise they provided while manufacturing all of the components needed for our test. We would also like to thank Joe Giliberto, who worked long into many nights to manufacture booms we needed before our flight test. References [1] Murphey, T. W., Jeon, S., Biskner, A., and Sanford, G., \u201cDeployable Booms and Antennas Using Bi-stable Tape-springs,\u201d 24th Annual AIAA/USU Conference on Small Satellites, Logan, UT, 2010, pp. SSC10\u2013X\u20136. [2] Lee, A. J., and Fernandez, J. M., \u201cMechanics of Bistable Two-Shelled Composite Booms,\u201d AIAA Spacecraft Structures Conference, 2018, pp. 1\u201324. doi:10.2514/6.2018-0938, URL https://arc.aiaa.org/doi/pdf/10.2514/6.2018-0938. [3] Fernandez, J. M., Rose, K., Younger, C. J., Dean, G. D., Warren, J. E., Stohlman, O. R., and Wilkie, W. K., \u201cNASA\u2019s advanced solar sail propulsion system for Low-cost deep space exploration and science missions that uses high performance rollable composite booms,\u201d 4th International Symposium on Solar Sailing, Japan Aerospace Exploration Agency, Tokyo, 2017, p. 11. URL https://ntrs.nasa.gov/search.jsp?R=20170001556. [4] Straubel, M., Block, J., Sinapius, M., and H\u00fchne, C., \u201cDeployable Composite Booms for Various Gossamer Space Structures,\u201d 52nd AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics and Materials Conference, AIAA, Denver, CO, 2011, pp. 1\u201311. doi:10.2514/6.2011-2023, URL http://arc.aiaa.org/doi/10.2514/6.2011-2023. [5] Hoskin, A., \u201cBlossoming of coiled deployable booms,\u201d 56th AIAA/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference, Vol. 1, No. January, 2015, pp. 1\u201310. doi:10.2514/6.2015-0207, URL https://arc.aiaa.org/doi/ abs/10.2514/6.2015-0207. D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58 [6] Mallikarachchi, H. M. Y. C., and Pellegrino, S., \u201cDeployment Dynamics of Ultrathin Composite Booms with TapeSpring Hinges,\u201d Journal of Spacecraft and Rockets, Vol. 51, No. 2, 2014, pp. 604\u2013613. doi:10.2514/1.A32401, URL http://arc.aiaa.org/doi/10.2514/1.A32401. [7] Russell, C. T. (ed.), The STEREO mission, Springer, Los Angeles, 2008. doi:10.1007/978-0-387-09649-0, URL http: //link.springer.com/content/pdf/10.1007/978-0-387-09649-0.pdf. [8] Fernandez, J. M., \u201cAdvanced Deployable Shell-Based Composite Booms For Small Satellite Structural Applications Including Solar Sails,\u201d International Symposium on Solar Sailing, Kyoto, Japan, 2017. URL https://ntrs.nasa.gov/search.jsp? R=20170001569. [9] Firth, J., and Pankow, M., \u201cAdvanced Dual-Pull Mechanism for Deployable Spacecraft Booms,\u201d Spacecraft Structures Conference, American Institute of Aeronautics and Astronautics, Kissimmee, FL, 2018, pp. 1\u20138. doi:10.2514/6.2018-0691, URL https://arc.aiaa.org/doi/10.2514/6.2018-0691. [10] Fernandez, J. M., Rose, G., Stohlman, O. R., Younger, C. J., Dean, G. D., Warren, J. E., Kang, J. H., Bryant, R. G., and Wilkie, K. W., \u201cAn Advanced Composites-Based Solar Sail System for Interplanetary Small Satellite Missions,\u201d 2018 AIAA Spacecraft Structures Conference, 2018, pp. 1\u201321. doi:10.2514/6.2018-1437, URL https://arc.aiaa.org/doi/10.2514/6.20181437. [11] Hibbeler, R. C., Engineering Mechanis: Dynamics, seventh ed., Pearson Prentice Hall, Upper Saddle River, New Jersey, 2007. [12] Weisstein, E. W., \u201cArchimedes\u2019 Spiral,\u201d , 2017. URL http://mathworld.wolfram.com/ArchimedesSpiral.html. [13] Fernandez, J. M., and Murphey, T. W., \u201cA Simple Test Method for Large Deformation Bending of Thin High Strain Composite Flexures,\u201d 2018 AIAA Spacecraft Structures Conference, 2018. doi:10.2514/6.2018-0942, URL https://arc.aiaa.org/ doi/10.2514/6.2018-0942. [14] Zero Gravity Corporation, \u201cZERO-G Research Program Package,\u201d , 2018. D ow nl oa de d by I O W A S T A T E U N IV E R SI T Y o n Ja nu ar y 11 , 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /6 .2 01 9- 12 58" + ] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure4-1.png", + "caption": "Fig. 4. First metamorphic limb with phase e2", + "texts": [ + " The metamorphic limb is synthesized by combining with the phase e1 in Fig. 3. Link a is fixed to base, and revolution joint of axis 2 is locked. Axis 1 rotate along with axis u1 direction. Axis 3 rotate along with w1 axis direction. Axis 4 and axis 5 are parallel to axis 3 with w1 direction. Then metamorphic limb takes on mobility four which can realize motion type submanifold T2(w1) \u00b7U(p1,w1,u1). Keeping rotational joint of axis 2 relaxed and rotational joint of axis 1 locked, rotate link c to axis 3 parallel with axis u1 direction in Fig. 4. Axis 2 rotate along with axis v1 direction, axis 4 and axis 5 are parallel to axis 3 with u1 direction. Then metamorphic limb satisfy form of R(q11,u1) \u00b7 R(q12,u1) \u00b7 R(q13,u1) and R(q14,v1) which generated by T2(u1) \u00b7U(q1,u1,v1) that have mobility four and can realize motion type such submanifold. 1) Type synthesis of metamorphic joint of second limb: T (3) \u00b7U(p2,w2,v2) and T (3) \u00b7U(q2,u2,v2) are submanifolds of second limb of motion type of 1R2T and 2R1T, respectively. Motion joint of second limb should satisfy one condition that transform freely between T (3) \u00b7U(p2,w2,v2) and T (3) \u00b7U(q2,u2,v2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002015_978-3-319-94346-6_6-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002015_978-3-319-94346-6_6-Figure6-1.png", + "caption": "Fig. 6. Schematic of route with multiple turnings.", + "texts": [ + " As in case of straight route takeoff and landing are under similar conditions but point 1 to 6 is in hover mode. Only difference is near point 3, Fig. 5, where UAV should change its direction for some degrees. Within this example in 40-degree angle is used. As result right side route (narrow angle) is shorter then left side route that express in difference of more than 3 kW to one or another side. Overhead lines outside short viewed sections as in previous examples are always complex with all possible combinations included (Fig. 6). In this route version will be more than two possible trajectories of flight (Fig. 7). Initially trajectories will be calculated in previous example with one turning on both sides. And in long distances such trajectories in sum will be almost the same. Left side flight also takes more power because of longer path, but difference was less then 2 kW on the same rout as single angled route. Also, angles are the same. It is not enough to consider of using left or right side of OHL to inspect it. To improve power consumption and make it more adaptive additional movements will be done" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000595_uksim.2009.59-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000595_uksim.2009.59-Figure1-1.png", + "caption": "Fig. 1 Description of the flexible manipulator system.", + "texts": [ + " Moreover, the FE method exhibits several advantages over the FD method [12]. However, in modelling of the manipulator using FE methods, the effects of beam\u2019s length have not been adequately addressed. The effect of length on the manipulator is important for modelling and control purposes, as successful implementation of a flexible manipulator control is contingent upon achieving acceptable uniform performance in the presence of length variations. The single-link flexible manipulator system considered in this work is shown in Fig. 1, where XoOYo and XOY represent the stationary and moving coordinates frames respectively, \u03c4 represents the applied torque at the hub. E, I, \u03c1, A, Ih and mp represent the Young modulus, area moment of inertia, mass density per unit volume, cross-sectional area, hub inertia and payload mass of the manipulator respectively. In this work, the motion of the manipulator is confined to XoOYo plane. Transverse shear and rotary inertia effects are neglected, since the manipulator is long and slender. Thus, the Bernoulli-Euler beam theory is allowed to be used to model the elastic behaviour of the manipulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001956_j.endm.2018.05.004-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001956_j.endm.2018.05.004-Figure1-1.png", + "caption": "Figure 1", + "texts": [ + " To interpolate between some a priori defined control points, a piecewise polar interpolant that approximate the polar trajectory can be expressed as a Hermite-type function [2,7] defined by r(\u03b8) = q\u2211 k=0 cik (\u03b8 \u2212 \u03b8i) k(4) where q = 3 (cubic approximation) is the order of the polynomial, ci0 = ri, ci1 = r\u0307i, c i 2 = 1 hi [(2r\u0307i + r\u0307i+1) + 3\u0394ri], c i 3 = 1 h2 i [r\u0307i + r\u0307i+1 \u2212 2\u0394ri], ri = r (\u03b8i), ri+1 = r (\u03b8i+1), hi = \u03b8i+1 \u2212 \u03b8i, \u0394ri = ri+1\u2212ri hi , and the derivatives at the endpoints are calculated using r\u0307 (\u03b8i) = dr(\u03b8i) d\u03b8 = r\u0307i and r\u0307 (\u03b8i+1) = dr(\u03b8i+1) d\u03b8 = r\u0307i+1 respectively. Considering the case of polar zenithal orthographic projection and related cylindrical coordinates (Fig. 1.c), trajectories representing the system dynamics can be generated using a combination of polar piecewise interpolants that approximate the polar trajectory, and Cartesian piecewise interpolants to approximate the trajectory height. Trajectory height is computed in the projective plane obtained by unfolding the cylinder with generators passing through the polar trajectory of the motion. Considering the case of polar gnomic or stereographic projection and related spherical coordinates (Fig. 1.a and Fig. 1.b), trajectories representing the system dynamics can be generated by a combination of polar piecewise interpolants over a \u201cspherical\u201d domain relatively similar with [6] by R(\u03bb, \u03c6) = S0(\u03c8) + S1(\u03c8)\u03c7+ S2(\u03c8)\u03c7 2 + S3(\u03c8)\u03c7 3(5) where \u03bb and \u03c6 represents the generalised coordinates related to the system model, S(\u03c8) = c0,0 + c1,0\u03c8 + c2,0\u03c8 2 + c3,0\u03c8 3, and and \u03c8 and \u03c7 are defined as in [6]. Therefore, adequate geometric considerations - such as manipulators geometry and related coordinate systems - used in the motion planning studies backed up by polar piecewise cubic interpolants allows easy calculation of kinematics variable and handling of the dynamical equations of motion [12] through the use of Lagrange\u2019s equations defined by d dt ( \u2202T \u2202q\u0307j ) \u2212 \u2202T \u2202qj = Qj(6) The kinetic energy T = and the acting contact and body forces Qi can be expressed as in [12] by Tj = 1 2 mjvCj \u00b7 vCj + 1 2 I\u0304j\u03c9j \u00b7 \u03c9j(7) and Qj = \u2202\u03c9 \u2202q\u0307j \u00b7T+ \u2202v \u2202q\u0307j \u00b7R(8) where v is the velocity, \u03c9 is the angular velocity, T is the applied torque, and R is the applied force" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003877_ijthi.2019070103-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003877_ijthi.2019070103-Figure5-1.png", + "caption": "Figure 5. Right avoidance: (a) front collision; (b) (stationary) collision", + "texts": [], + "surrounding_texts": [ + "In\ufeffthe\ufefftheory\ufeffof\ufeffcomputer\ufeffsimulation,\ufeffthe\ufeffpositions\ufeffof\ufeffvirtual\ufeffmembers\ufeffare\ufeffconfirmed\ufeffby\ufeffcoordinates.\ufeff A\ufefftwo-dimensional\ufeffplane\ufeffis\ufeffconstructed\ufeffto\ufeffsimulate\ufeffthe\ufeffincident\ufeffsite\ufeffin\ufeffthis\ufeffpaper.\ufeffEvery\ufeffposition\ufeffon\ufeff the\ufeffplane\ufeffhas\ufeffa\ufeffspecific\ufeffcoordinate\ufeff ( , )x y ,\ufeffcorresponding\ufeffto\ufeffeach\ufefflocus\ufeffof\ufeffevery\ufeffvirtual\ufeffplayer.\ufeffThe\ufeff coordinate\ufeffvariety\ufeffindicates\ufeffchanging\ufeffof\ufeffthe\ufeffplayer\u2019s\ufeffposition,\ufeffand\ufeffthe\ufeffdifference\ufeffof\ufeffthe\ufeffcoordinate\ufeff values\ufeffrepresents\ufeffthe\ufeffdistance\ufeffbetween\ufefftwo\ufeffplayers.\ufeffWhen\ufeffthe\ufeffformation\ufeffchanges,\ufeffthe\ufeffnext\ufeffstep\ufeffforward\ufeff of\ufeffthe\ufeffmembers\ufeffcan\ufeffbe\ufeffobtained\ufeffon\ufeffthe\ufeffbasis\ufeffof\ufeffthe\ufefforiginal\ufeffand\ufeffgoal\ufeffcoordinates. This\ufeffpaper\ufeffregards\ufeffthe\ufefflocus\ufeffwhere\ufeffthe\ufeffgroup\ufeffformation\ufefflocates\ufeffas\ufeffa\ufefftwo-dimensional\ufeffplane.\ufeffThe\ufeff moving\ufeffdirections\ufeffof\ufeffthe\ufeffvirtual\ufeffplayers\ufeffare\ufeffindicated\ufeffby\ufeffthe\ufeffincrements\ufeff \u2206x \ufeffand\ufeff \u2206y \ufeffin\ufeffx-axis\ufeffand\ufeff y-axis\ufeff coordinate\ufeff directions.\ufeff The\ufeff values\ufeff of\ufeff \u2206x \ufeff and\ufeff \u2206y \ufeff are\ufeff continuously\ufeff recorded\ufeff at\ufeff a\ufeff certain\ufeff frequency,\ufeffbase\ufeffon\ufeffwhich\ufeffthe\ufeffpath\ufeffof\ufeffthe\ufeffvirtual\ufeffplayers\ufeffcould\ufeffbe\ufeffknown.\ufeffUsing\ufeffthe\ufeffpath\ufeffconversion\ufeff algorithm,\ufeffthe\ufeffentire\ufeffconversion\ufeffprocedure\ufeffcould\ufeffbe\ufeffsimulated\ufeffto\ufeffdisplay\ufeffthe\ufeffdetailed\ufefftrajectory\ufeffof\ufeff every\ufeffplayer\u2019s\ufeffpath\ufeffchange. Every\ufeffgroup\ufeffis\ufeffrequired\ufeffto\ufeffkeep\ufeffaggregated\ufeffand\ufeffsustainable\ufeffafter\ufefftransforming\ufeffformation.\ufeffThereby,\ufeff the\ufefftransformed\ufeffformation\ufeffis\ufeffdivided\ufeffinto\ufeffpolygon\ufeffregions\ufeffbased\ufeffon\ufeffthe\ufeffquantity\ufeffof\ufeffleaders.\ufeffFirst,\ufeff the\ufeffclosest\ufeffpoint\ufeffto\ufeffthe\ufeffregional\ufeffcenter\ufeffis\ufeffset\ufeffas\ufeffthe\ufeffposition\ufeffof\ufeffleader,\ufeffand\ufeffthen,\ufeffbased\ufeffon\ufeffgreedy\ufeff algorithm,\ufeffleader\ufeffare\ufeffassigned\ufeffto\ufeffdifferent\ufeffregions\ufeffto\ufefffeedback\ufeffreal-time\ufefflocation,\ufefffinally,\ufeffcollision\ufeff will\ufeffbe\ufeffpredicted,\ufeffdetected\ufeffand\ufeffavoided\ufeffin\ufeffthe\ufeffmovement. Formation control with Greedy Algorithm Based on conditional Formation Feedback Aiming\ufeffat\ufeffcoordinating\ufeffthe\ufeffactions\ufeffbetween\ufeffobstacle\ufeffavoidance\ufeffand\ufeffformation\ufeffkeeping,\ufeffwhile\ufeffone\ufeff player\ufeff is\ufeff avoiding\ufeff obstacles,\ufeff the\ufeff others\ufeff must\ufeff keep\ufeff the\ufeff formation\ufeff operating\ufeff orderly\ufeff based\ufeff on\ufeff the\ufeff greedy\ufeffalgorithm\ufeff(Pan\ufeff&\ufeffRuiz,\ufeff2014).\ufeffMembers\ufeffdon\u2019t\ufeffhave\ufeffany\ufeffrelations\ufeffwhile\ufeffmaintaining\ufeffformation\ufeff because\ufeff of\ufeff the\ufeff parallel\ufeff structure,\ufeff which\ufeff makes\ufeff formation\ufeff control\ufeff more\ufeff flexible\ufeff and\ufeff stable\ufeff (Xin,\ufeff Daqi\ufeff&\ufeffYangyang,\ufeff2017).\ufeffIn\ufeffthe\ufeffmovement,\ufeffFormation\ufefffeedback\ufeffis\ufeffconducted\ufeffin\ufefforder\ufeffto\ufeffresolve\ufeffthe\ufeff issue\ufeffthat\ufeffsome\ufefffollower\ufeffmay\ufeffbe\ufeffout\ufeffof\ufeffthe\ufeffrange\ufeffof\ufeffcommunication\ufeffbecause\ufeffof\ufeffobstacle\ufeffavoidance.\ufeff Consequently,\ufeffit\ufeffprobably\ufeffleads\ufeffto\ufeffthe\ufeffobstruction\ufeffin\ufeffthe\ufeffwhole\ufeffmovement,\ufeffand\ufeffsystem\ufeffwill\ufeffstay\ufeffthe\ufeff locked\ufeffstate\ufeffwhile\ufeffmultiple\ufefffollowers\ufeffare\ufeffstuck\ufeffin\ufeffbarrier\ufeffsimultaneously.\ufeffThereby,\ufeffthe\ufeffintroduction\ufeff of\ufeffconditional\ufefffeedback\ufeff(Singh\ufeff&\ufeffAhmed,\ufeff2014)\ufeffis\ufeffto\ufeffsolve\ufeffthis\ufeffproblem,\ufeffas\ufeffshown\ufeffin\ufeffFigure\ufeff3. Simulate\ufeffa\ufefftwo-dimensional\ufeffplane\ufeffwith\ufeffn\ufeffpoints\ufeffrepresenting\ufeffthe\ufeffleaders,\ufeffand\ufeffimport\ufeffa\ufeffmatrix\ufeff whose\ufeffitems\ufeffrepresent\ufeffthe\ufeffdistances\ufeffbetween\ufeffthe\ufefforiginal\ufefflocation\ufeffand\ufeffthe\ufeffgoal\ufeffof\ufeffleaders.\ufeffIn\ufeffthis\ufeffway,\ufeff the\ufeffproblem\ufeffof\ufeffcalculating\ufeffthe\ufeffshortest\ufeffroute\ufeffwill\ufeffbe\ufefftransmuted\ufeffinto\ufeffan\ufeffissue\ufeffof\ufeffsearching\ufeffthe\ufeffminimum\ufeff value\ufeffin\ufeffthe\ufeffmatrix,\ufeffwhich\ufeffis\ufeffmarked\ufeffLn n a ij\u00d7 = ( ) .\ufeffWhere\ufeff i \ufeffrepresents\ufeffthe\ufeffnth\ufeffleader\ufeffof\ufefforiginal\ufeff position;\ufeff j \ufeffstands\ufefffor\ufeffthe\ufefflocation\ufeffof\ufeffthe\ufeffi-th\ufeffleader\ufeffin\ufeffthe\ufeffobject\ufeffformation;\ufeffa ij \ufeffis\ufeffthe\ufeffdistance\ufefffrom\ufeff the\ufefforiginal\ufefflocation\ufeffto\ufeffthe\ufeffgoal\ufeffof\ufeffthe\ufeffi-th\ufeffleader.\ufeffWhile\ufefftransformation,\ufeffall\ufeffthe\ufeffvirtual\ufeffplayers\ufeffmove\ufeff orderly\ufeffand\ufeffthen\ufeffthe\ufeffwhole\ufeffformation\ufeffcould\ufeffbe\ufefftransformed\ufefffrom\ufeffinitial\ufeffshape\ufeffto\ufefftarget\ufeffshape.\ufeffExcept\ufeff Algorithm 1. Example of a greedy algorithm Void\ufeffContainerLoading(int\ufeffx[],float\ufeffw[],float\ufeffc,int\ufeffn)\ufeff {\ufeff//x[i]=1\ufeffWhen\ufeffand\ufeffonly\ufeffwhen\ufeffthe\ufeffcontainer-i\ufeffis\ufeffloaded\ufeffon\ufeffthe\ufeffweight,\ufeffsorted\ufeffby\ufeffindirect\ufeffaddressing\ufeffmode\ufeff int*t=new\ufeffint[n+1];\ufeff//table\ufeffn\ufeffis\ufeffindirect\ufeffaddressing\ufeff IndirectSort(w,t,n);\ufeff//now,\ufeffw[t(i)]\ufeffw[t(i+1)],1\ufeffi\ufeffn\ufeff \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0for(int\ufeffi=1;i<=n;i++)//initialize\ufeffx\ufeff \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0X[i]=0;\ufeff for(i=1;i<=n)\ufeff&&\ufeffw[t(i)]<=c;i++{\ufeff//select\ufeffitems\ufeffby\ufeffweight\ufeff x[t(i)]=1;\ufeff c=w[t(i)];\ufeff }\ufeff//residual\ufeffcapacity\ufeff delete\ufefft[]; for\ufeffthe\ufeffmovement\ufeffroutes,\ufeffThe\ufeffposition\ufeffof\ufeffeach\ufeffmember\ufeffis\ufefffixed\ufeffboth\ufeffin\ufefforiginal\ufeffand\ufeffgoal\ufeffformation,\ufeff suppose\ufeffthe\ufeffdistance\ufeffis\ufeffa\ufeffroute\ufeffbetween\ufeffthe\ufeffstarting\ufeffpoint\ufeffand\ufeffthe\ufefffoothold,\ufeffand\ufeffthe\ufeffpriority\ufeffalgorithm\ufeff aims\ufeffto\ufefflook\ufefffor\ufeffan\ufeffideal\ufeffroute\ufeffwhich\ufeffis\ufeffthe\ufeffshortest\ufeffone\ufeffwithout\ufeffany\ufeffconflict\ufeff(Bj\u00f8rn\u00f8y,\ufeffBassett\ufeff&\ufeff Ucar,\ufeff2016;\ufeffMa,\ufeffTang\ufeff&\ufeffWang,\ufeff2016).\ufeffA\ufeffmember\ufeffcan\ufeffbe\ufeffaccomplished\ufeffthe\ufefftask\ufeffby\ufeffrunning\ufeffforward\ufeff from\ufeffthe\ufeffinitial\ufeffpoint\ufeff to\ufeffthe\ufeffgoal\ufeffformation.\ufeffAccording\ufeffto\ufeffthe\ufeffdistance\ufeffcalculated\ufefffrom\ufeffthe\ufeffinitial\ufeff location\ufeffto\ufeffthe\ufeffgoal,\ufeffa\ufeffdistance\ufeffmatrix\ufeffof\ufeffn n\u00d7 \ufeffis\ufeffobtained. 26 15 37 20 12 11 14 21 19 12 9 13 8 24 11 23 \u2192 26 15 11 37 20 21 9 13 23 \ufeff (6) So\ufeffthe\ufeffchange\ufeffin\ufeffthe\ufeffthis\ufeffmatrix\ufeffindicates\ufeffthat\ufeffthe\ufeffshortest\ufeffroute\ufeffis\ufeffa 33 8= . collision Prediction The\ufeffpositions\ufeffof\ufeffthe\ufeffwhole\ufeffformation\ufeffmembers\ufeffin\ufeffthe\ufeffgoal\ufeffformation\ufeffare\ufeffcalculated\ufeffusing\ufeffconditional\ufeff feedback\ufeff(Andrade\ufeff&\ufeffSantos,\ufeff2017;\ufeffBarron,\ufeffCohen\ufeff&\ufeffDahmen,\ufeff2008).\ufeffHowever,\ufeffin\ufeffthe\ufeffwhole\ufeffprocess\ufeff of\ufeffthe\ufeffformation\ufefftransformation,\ufeffit\ufeffis\ufeffinevitable\ufeffthat\ufefftwo\ufeffor\ufeffmore\ufeffmembers\ufeffcome\ufeffacross\ufeffsimultaneously,\ufeff and\ufeffit\ufeffis\ufeffcollision.\ufeffIf\ufeffthe\ufeffproblem\ufeffhad\ufeffnot\ufeffbeen\ufeffhandled\ufeffproperly,\ufeffthe\ufeffwhole\ufeffformation\ufeffwould\ufeffhave\ufeff be\ufeffat\ufeffa\ufeffstandstill\ufeff(Niewiadomska-Szynkiewicz\ufeff&\ufeffSikora,\ufeff2011).\ufeffAccordingly,\ufeffcollision\ufeffneeds\ufeffto\ufeffbe\ufeff predicted\ufeffthrough\ufeffthe\ufeffwhole\ufeffprocess. In\ufeffthis\ufeffstudy,\ufeffx \ufeffstands\ufefffor\ufeffa\ufeffformation\ufeffmember,\ufeffL x( ) stands\ufefffor\ufeffthe\ufeffpresent\ufefflocation\ufeffof\ufeffx ,\ufeffV x( ) \ufeff is\ufeffthe\ufeffvelocity\ufeffof\ufeff x ,\ufeffand\ufeffW x( ) \ufeffis\ufeffthe\ufeffwidth\ufeffof\ufeffregion,\ufeffwhich\ufeffis\ufeffa\ufeffconstant\ufeffin\ufeffthis\ufeffstudy\ufeff.\ufeffSuppose\ufeff that\ufeff2\ufeffadjacent\ufeffplayers\ufeffare\ufeffa \ufeffand\ufeffb ,\ufeffand\ufeffL L a L b r = ( ) ( )- ,\ufeffV V a V b r = ( ) ( )- ,\ufeffwhere\ufeffL r \ufeffstands\ufefffor\ufeffthe\ufeff relative\ufeffposition\ufeffof\ufeffa \ufeffand\ufeffb ,\ufeffand\ufeffV r \ufeffthe\ufeffrelative\ufeffvelocity\ufeffof\ufeffa \ufeffand\ufeffb ,\ufeffso\ufeffa\ufeffprobable\ufeffcollision\ufeffbetween\ufeff a \ufeffand\ufeffb \ufeffshould\ufeffsatisfy: L L V t V t W a W b r r r r 2 2 2 22+ \u00d7 \u00d7 \u00d7 + \u00d7 = + +( )( ) ( ) \u03b5 \ufeff (7) Where\ufeff\u03b5 \ufeffis\ufeffthe\ufeffsafe\ufeffdistance\ufeffof\ufeffa \ufeffand\ufeffb .\ufeffThe\ufeffexistence\ufeffof\ufeffunique\ufeffsolution\ufeffor\ufeffunbounded\ufeffsolution\ufeff predicts\ufeffthat\ufeffa\ufeffwould\ufeffnot\ufeffcollide\ufeffwith\ufeffb.\ufeffIf\ufefftwo\ufeffsolutions\ufeffof\ufefft 1 \ufeffand\ufefft 2 t t 1 2 <( ) \ufeffexist,\ufeffit\ufeffindicates\ufeffthat\ufeff there\ufeffwill\ufeffbe\ufeffa\ufeffcollision\ufeffinstantly\ufeff(so\ufeffcollision\ufeffavoidance\ufeffis\ufeffessential\ufeffin\ufeffthis\ufeffcase)\ufeff;\ufeffif\ufefft 2 0< ,\ufeffa\ufeffwould\ufeff collide\ufeff with\ufeff b\ufeff after\ufeff t 1 .\ufeff The\ufeff time\ufeff of\ufeff collision\ufeff could\ufeff be\ufeff indicated\ufeff uniformly\ufeff with\ufeff t p ,\ufeff then\ufeff L L a V a t a p = + \u00d7( ) ( ) ,\ufeffL L b V a t b p = + \u00d7( ) ( ) ,\ufeffwhere\ufeffL a \ufeffand\ufeffL b \ufeffstands\ufefffor\ufeffthe\ufeffpositions\ufeffof\ufeffa \ufeffand\ufeff b \ufeffafter\ufeffthe\ufefftime\ufeff t p \ufeffrespectively,\ufeffand\ufeffthen: 1.\ufeff\ufeff L L V a a b \u2212( )\u00d7 <( ) 0 \ufeffrepresents\ufeffrear\ufeffcollision, 2.\ufeff\ufeff L L V a a b \u2212( )\u00d7 >( ) 0 \ufeffand\ufeffV a V b( ) ( )\u00d7 < 0 \ufeffrepresent\ufefffront\ufeffcollision, 3.\ufeff\ufeff L L V a a b \u2212( )\u00d7 >( ) 0 \ufeffand\ufeffV a V b( ) ( )\u00d7 \u2265 0 \ufeffrepresent\ufeffrear\ufeffcollision, 4.\ufeff\ufeff V b( ) = 0 \ufeffrepresents\ufeffstationary\ufeffcollision. The\ufefflocalized\ufeffcollision\ufeffavoidance\ufeffalgorithm\ufeffis\ufeffdeveloped\ufeffbased\ufeffon\ufeffthese\ufefffour\ufefftypes. collision Avoidance Collision\ufeffavoidance\ufeff(Ajorlou,\ufeffAsadi\ufeff&\ufeffAghdam,\ufeff2015;\ufeffLozano-P\u00e9rez\ufeff&\ufeffWesley,\ufeff1979)\ufeffin\ufeffthis\ufeffpaper\ufeff is\ufeffprimarily\ufeffconducted\ufeffby\ufeffthe\ufeffchange\ufeffof\ufeffvelocity\u2019s\ufeffmagnitude\ufeffand\ufeffdirection.\ufeffSo\ufeffthe\ufeffavoidance\ufeffcan\ufeffbe\ufeff achieved\ufeffwith\ufeffthe\ufeffchange\ufeffof\ufeffmagnitude\ufefflike\ufeffacceleration\ufeffor\ufeffdeceleration\ufeffand\ufeffdirection\ufefflike\ufeffleft\ufeffor\ufeff right,\ufeffsimultaneously\ufeffor\ufeffnot.\ufeffThe\ufeffacceleration\ufeffof\ufeffa \ufeffis\ufeffequal\ufeffto\ufeffthe\ufeffdeceleration\ufeffof\ufeffb ,\ufeffso\ufeffthe\ufeffdeceleration\ufeff rule\ufeffcould\ufeffbe\ufeffrealized\ufeffby\ufeffthe\ufeffacceleration\u2019s.\ufeffThe\ufeffdiagrams\ufeffof\ufeffthe\ufeffleft,\ufeffthe\ufeffright\ufeffand\ufeffthe\ufeffvelocity-vary\ufeff avoidance\ufeffare\ufeffrespectively\ufeffdisplayed\ufeffin\ufeffFigure\ufeff4,\ufeffFigure\ufeff5\ufeffand\ufeffFigure\ufeff6. Formation Initial simulation The\ufeffsimulation\ufeffis\ufeffimplemented\ufeffcombining\ufeffthe\ufeffgreedy\ufeffalgorithm\ufeffand\ufeffconditional\ufefffeedback. 1.\ufeff\ufeff Simulation Environment:\ufeffLength\ufeffl\ufeff=\ufeff200,\ufeffwidth\ufeffw\ufeff=\ufeff200,\ufeffgrid\ufeffsize\ufeffc\ufeff=\ufeff1;\ufeffgrid\ufeffdata\ufeffX\ufeffaxis\ufeff direction\ufeff N x =1/ e ;\ufeff grid\ufeff data\ufeff Y\ufeff axis\ufeff direction\ufeff N w c y = / ;\ufeff X\ufeff axis\ufeff direction\ufeff target\ufeff value\ufeff des N x _ / )x=round( 2 10\u2212 ;\ufeffY\ufeffaxis\ufeffdirection\ufefftarget\ufeffvalue\ufeffdes y _y=N \u221220 . 2.\ufeff\ufeff Two-dimensional\ufeffnetwork\ufeffenvironment\ufeffconstruction:Environment\ufefffunction\ufeffA initialize N N x y = ( , ) \ufeff and\ufeff A ones N N x y = \u00d72 ( , ) .\ufeffThe\ufefffunction\ufeffsets\ufeffthe\ufeffrandom\ufeffposition\ufeffand\ufefftarget\ufeffposition\ufeffof\ufeffthe\ufeff initial\ufeffteam.\ufeffFor\ufeffexample,\ufeffthe\ufefffour\ufeffplayers\u2019\ufeffpositions\ufeffin\ufeffthe\ufefffollowing\ufeffsimulation\ufeffdiagram\ufeffare\ufeff marked\ufeffas\ufefffollowed:\ufeffA(90,50)\ufeff=\ufeff1;\ufeffA(30,30)\ufeff=\ufeff1;\ufeff(50,40)\ufeff=\ufeff1;\ufeffA(100,20)\ufeff=\ufeff1;\ufeffwhere\ufeffthe\ufeffred\ufeff pattern\ufeffrepresents\ufeffthe\ufeffmembers. 3.\ufeff\ufeff The Determination of the Leader:\ufeffFirstly,\ufeffthe\ufeffmembers\ufeffin\ufeffthe\ufeffenvironment\ufeffare\ufeffsearched,\ufeffand\ufeff the\ufeffcoordinates\ufeffand\ufefftheir\ufeffnumbers\ufeffare\ufeffwritten\ufeffdown,\ufeffand\ufeffthen\ufeffthe\ufeffdistance\ufeffbetween\ufeffthe\ufeffmember\ufeff first\ufeffsearched\ufeffand\ufeffthe\ufefftarget\ufeffpoint\ufeff(task\ufefffinal\ufefftarget\ufeffpoint)\ufeffis\ufeffcalculated\ufeffas\ufeffa\ufeffbasis\ufefffor\ufeffcomparison.\ufeff Secondly,\ufeffthe\ufeffdistance\ufeffbetween\ufeffthe\ufefftarget\ufeffpoint\ufeffand\ufeffother\ufeffmembers\ufeffis\ufeffcalculated,\ufeffand\ufeffthen\ufeffthe\ufeff minimum\ufeffdistance\ufeffvalue\ufeffis\ufeffobtained.\ufeffFinally,\ufeffthe\ufeffnearest\ufeffmember\ufefffrom\ufeffthe\ufefftarget\ufeffpoint\ufeffis\ufeffset\ufeffas\ufeff the\ufeffleader. 4.\ufeff\ufeff The Formation of Team Members:\ufeffTo\ufefftake\ufeffa\ufeffdiamond\ufefffor\ufeffexample,\ufefffirst\ufeffdetermine\ufeffall\ufeff the\ufeff geometric\ufeffform\ufeffposition,\ufeffand\ufeffthe\ufeffpoint\ufeffcode\ufeffis\ufeffas\ufeffbelow: position_x\ufeff=\ufeff[leader_xleader_x-40leader_x-40leader_x-80];\ufeffposition_y\ufeff=\ufeff[leader_yleader_y10leader_y+10leader_y].\ufeff Determine\ufeffthe\ufeffgeometric\ufefftarget\ufeffposition\ufeffof\ufeffthe\ufefffour\ufeffmembers.\ufeffFirstly\ufeffthe\ufeffdistance\ufefffrom\ufeffthe\ufeffi-th\ufeff follower\ufeffto\ufeffthe\ufeffleader\ufeffis\ufeffcalculated,\ufeffsecondly\ufefffrom\ufeffthe\ufeffi-th\ufefffollower\ufeffto\ufeffits\ufeffgoal\ufeffpoint,\ufeffultimately\ufeffthe\ufeff closest\ufeffposition\ufeffapart\ufefffrom\ufeffthe\ufeffi-th\ufefffollower\ufeffis\ufeffobtained,\ufeffand\ufeffits\ufeffobject\ufeffposition\ufeffin\ufeffthe\ufeffdiamond\ufeffis\ufeff found.\ufeffAnd\ufeffthe\ufeffcode\ufeffis\ufeffas\ufeffbelow: < (b) a arrives the collision position far before for i=1:N if (r_x(i)==leader_x)&&(r_y(i)==leader_y) r_px(i)=leader_x; r_py(i)=leader_y; position_x(i)=1000; leader=i; end end for i=1:N if i==leader else r_px(i)=position_x(l); r_py(i)=position_y(l); r_p(i)=sqrt((r_x(i)-position_x(l))^2+(r_y(i)position_y(l))^2); for j=2:length(position_x) l=sqrt((r_x(i)-position_x(j))^2+(r_y(i)-position_y(j))^2); if l \u2264( )4 3 4&& ,\ufeffthen\ufeffthe\ufeffmovement\ufeffprocess\ufeffis\ufeffas\ufefffollows:\ufeffthe\ufefffollower\ufeffmoves\ufeffone\ufeff grid\ufeffhorizontally\ufeffwhen\ufeffA r x r y_ , _+( ) ==1 1 ,\ufeffotherwise\ufeffmoves\ufeffone\ufeffgrid\ufeffvertically. grid\ufeffhorizontally\ufeffwhen\ufeffA r x r y_ , _+( ) ==1 1 ,\ufeffotherwise\ufeffretreats\ufeffone\ufeffgrid\ufeffvertically. Figure\ufeff7\ufeffindicates\ufeffthat\ufeffthe\ufeffformation\ufeffcan\ufeffbe\ufeffwell\ufeffmaintained\ufeffintroducing\ufeffformation\ufeffcontrol\ufeffbased\ufeff on\ufeffgreedy\ufeffalgorithm\ufeffwith\ufeffconditional\ufefffeedback\ufeffwhen\ufeffavoiding\ufeffobstacles. sIMULATIoN eXPerIMeNT ANd resULT ANALysIs dynamic Model Based on Leader-Follower Algorithm According\ufeffto\ufeffspecific\ufeffmission\ufeffrequirements,\ufeffunits\ufeffin\ufeffemergencies\ufeffprocess\ufeffcan\ufeffbe\ufeffable\ufeffto\ufeffchange\ufeffand\ufeff maintain\ufeffspecific\ufeffgeometric\ufeffshapes,\ufeffand\ufeffrealize\ufeffformation\ufeffand\ufefftransformation\ufeffin\ufeffreal\ufefftime\ufeffaccording\ufeff to\ufeffchanges\ufeffin\ufeffthe\ufeffenvironment\ufeffto\ufeffbetter\ufeffcomplete\ufeffthe\ufeffassignment\ufeffPeng,\ufeff(Dimarogonas,\ufeffTsiotras\ufeff&\ufeff Kyriakopoulos,\ufeff2009;\ufeffWen\ufeff&\ufeffRahmani,\ufeff2013).\ufeffFor\ufeffpractical\ufeffproblems,\ufeffthe\ufeffformation\ufeffcan\ufeffbe\ufeffvaried\ufeff according\ufeffto\ufeffthe\ufeffenvironmental\ufefffactors\ufeffand\ufeffthe\ufeffdifficulty\ufeffof\ufeffdifferent\ufefftasks.\ufeffIn\ufeffdealing\ufeffwith\ufeffemergencies,\ufeff the\ufeffformation\ufeffof\ufeffthe\ufefftransformation\ufeffoften\ufeffuse\ufeffa\ufeffdiamond-shaped\ufeffdeformation\ufeffH-shaped\ufeffformation,\ufeff horizontal\ufeff team\ufeffformation\ufeffdiamond-shaped\ufeffformation,\ufeffdiamond-shaped\ufeffteam\ufeffdeformation\ufeffand\ufeffso\ufeff on(Lin,\ufeffHwang\ufeff&\ufeffWang,\ufeff2014;\ufeffNoguchi,\ufeffOhtaki\ufeff&\ufeffKamada,\ufeff2016).\ufeffIt\ufeffshows\ufeffthe\ufeffvertical,\ufefftriangular,\ufeff diamond\ufeffand\ufeffwedge\ufeffformation\ufefffrom\ufeffleft\ufeffto\ufeffright\ufeffin\ufeffFigure\ufeff8. To\ufeff confirm\ufeff the\ufeff practicability\ufeff of\ufeff this\ufeff approach,\ufeff this\ufeff study\ufeff validates\ufeff the\ufeff formation\ufeff from\ufeff the\ufeff horizontal\ufeffteam\ufeffinto\ufeffa\ufeffdiamond\ufeffand\ufefftrapezoidal,\ufeffand\ufeffeffect\ufeffof\ufefftransformation\ufeffis\ufeffindicated\ufeffin\ufeffFigure\ufeff9. The\ufeffexperiment\ufeffis\ufeffimplemented\ufeffwith\ufeffMatlab.\ufeffFirstly,\ufeff90\ufeffpositions\ufeffof\ufeffthe\ufefforiginal\ufeffformation\ufeffare\ufeff preset\ufeffand\ufeffassigned\ufeffto\ufeff10\ufeffgroups,\ufeffand\ufeffthe\ufeffcenter\ufeffof\ufeffevery\ufeffgroup\ufeffis\ufeffthe\ufeffleader.\ufeffThen,\ufeff90\ufeffpositions\ufeff of\ufeffgoal\ufeffformation\ufeffare\ufeffprovided\ufeffand\ufeffsplit\ufeffinto\ufeff10\ufeffzones\ufeffby\ufeffpolygons,\ufeffwhere\ufeffthe\ufeffnearest\ufeffpoint\ufeffto\ufeffthe\ufeff regional\ufeffcenter\ufeffis\ufeffthe\ufeffposition\ufeffof\ufeffleader.\ufeffBased\ufeffon\ufeffthe\ufeffgreedy\ufeffalgorithm,\ufeffthe\ufeffmapping\ufeffcoordinates\ufeffof\ufeff the\ufeffleader\ufeffand\ufeffmembers\ufefffrom\ufeffinitial\ufeffformation\ufeffto\ufeffthe\ufeffgoal\ufeffis\ufeffdrawn.\ufeffFinally,\ufeffcollision\ufeffavoidance\ufeffis\ufeff implemented\ufeffbased\ufeffon\ufeffthe\ufeffcontext\ufeffavoidance\ufeffrules\ufeff(as\ufeffshown\ufeffin\ufeffconversion\ufeff2). Figure\ufeff9\ufeffshows\ufeffthat\ufeffformations\ufeffstart\ufeffto\ufeffaggregate\ufeffat\ufeffconversion\ufeff1\ufeffwhere\ufeffthe\ufeffleaders\ufeffemerge;\ufeffand\ufeff the\ufeffgoal\ufeffformation\ufeffappears\ufeffin\ufeffconversion\ufeff2.\ufeffDuring\ufeffthe\ufeffwhole\ufefftransformation\ufeffprocess,\ufeffthere\ufeffis\ufeffno\ufeff collision\ufeffbecause\ufeffof\ufeffgaps\ufeffbetween\ufeffpoints.\ufeffIn\ufeffview\ufeffof\ufefftransformation,\ufeffthe\ufefffeasible\ufeffpath\ufeffis\ufeffachieved\ufeffby\ufeff greedy\ufeffalgorithm,\ufeffyet\ufeffnot\ufeffthe\ufeffoptimum." + ] + }, + { + "image_filename": "designv11_92_0000213_s11665-008-9314-5-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000213_s11665-008-9314-5-Figure1-1.png", + "caption": "Fig. 1 Outline showing the system of fixation and the direction of the forces to which it is submitted the spring steel", + "texts": [ + " The combination of adding alloying elements to steels and more ambitious production processes have proved to be a quite viable way in obtaining the required results. This tendency has also been observed in the production of suspension systems of automotives vehicles, which are constituted of leaf springs kept together using U-bolts. The bolt aims to maintain the leaf springs attached to the axis of the vehicle, forming a three components solid set: the axis, the leaf springs, and the supporting plate (see Fig. 1). These parts are combined in the vehicles suspension system, forming a complex union with the U-bolt, and each of them presents its own effect on the performance of the whole system. They work under action and reaction forces, always in the opposite direction than that of the springs; the bolts themselves support all the tensile stresses generated by the suspension during the working cycles. The suspension once is set, will be dynamically loaded and a minimum force of turning moment should previously be applied or the U-bolt will fail by fatigue (Ref 1, 2). Figure 1 shows the suspension system submitted to a force F originated from the maximum stress to which the spring may be submitted during work (Ref 3). Thus, the bolts support the maximum tensile stress. Geometrically, the bolts are square or round as showed in Fig. 2(a) and (b), respectively. Bolts are usually manufactured employing SAE 4140 steel in the quenched and tempered condition to obtain a tempered martensite microstructure to ensure the level of the necessary mechanical resistance for this type of component" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001996_icasi.2018.8394258-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001996_icasi.2018.8394258-Figure1-1.png", + "caption": "Fig. 1 A small underwater vehicle with distributed CMGs.", + "texts": [ + " That is, practical effects induced by the attitude control will be easily affected by the underwater environment with complex flow field disturbances. Motivated by the above problems and more diverse demands, internal attitude control devices are becoming a cause for concern. In order to take full advantages of the internal space of an underwater vehicle, dispersed-installed gyro units can be applied in the vehicle to control the attitude 28 ISBN 978-1-5386-4342-6 and propeller is used to change the position. For instance, three gyro units are distributed inside of the underwater vehicle body shell (see Fig.1), which can form a triangular pyramid CMGs configuration. For this special underwater vehicle, the distributed CMGs can generate multi-direction torque to control its attitude at a low speed or even zero speed, and the attitude under a high speed can be controlled by the propeller combined with the internal CMG system to increase its own agility and maneuverability. That is, two ways of the attitude control program can be fulfilled by the controller, which is installed inside the body to contact the power and gyro components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003978_sii.2019.8700356-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003978_sii.2019.8700356-Figure3-1.png", + "caption": "Fig. 3. Inertial Frame and Body-Fixed Frame of UAV", + "texts": [ + " Then, as shown in [4], the state equation of the UAVs is given by \u03be\u0307 = v, mv\u0307 = mge3\u2212 f Re3, R\u0307 = R\u2126\u0302, J\u2126\u0307 = M\u2212\u2126\u00d7 J\u2126, (7) where m is the total mass of the UAV, J is the inertia matrix with respect to the body-fixed frame, g is the acceleration of gravity, e3 = (0,0,1)T , f \u2208R is the thrust, M \u2208R3 is the total moment in the body fixed frame, and \u2126\u0302 denotes the matrix satisfying \u2126\u0302y = \u2126\u00d7 y for all y \u2208 R3. In this paper, we develop the coverage control using multiple UAVs based on the coverage control with the integrators. Fig. 3 shows the relation between the inertial frame and the body-fixed frame where the inertial frame and the body-fixed frame are denoted by {A} and {B}, respectively. To this end, we need to determine the thrust f and the moment M in the state equation of the UAV (7) for the coverage control. We first discuss the coverage control method with the integrators for moving regions. The procedure to control the method for moving regions are as follows. In [7], a distributed controller for the coverage control is shown" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000619_6.2008-493-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000619_6.2008-493-Figure3-1.png", + "caption": "Figure 3: System layout of the Hyperion", + "texts": [ + " The Hyperion must be able to fly autonomously and be remote controlled and for this it has to have advanced avionic systems on board, which are mostly commercial of the shelf (COTS) products and are located in modules in the front of the fuselage. Fuel is located in the wings and in several tanks in the fuselage and the landing gear is also located in the fuselage. An overview of these systems can be found in . Other important aspects are discussed below. Parameter Value Aspect ratio [-] 4 Wing area [m2] 3.78 Wingspan [m] 3.89 Mean wing chord [m] 1.04 Quarter chord sweep [\u00b0] 40 Taper ratio [-] 0.16 Thickness/chord ratio [-] 0.1 Fuselage length [m] 2.65 Table 2: Parameter values of the Hyperion Figure 3 The advantage of a rocket assisted take-off (RATO) is a reduced time needed to get the IUAV airborne. This will not necessarily lead to a greater intercept distance, as is shown in chapter VII. The advantage of a parachute recovery is the flexibility in getting the IUAV safely onto the ground at any location, not only airfields, especially in case of an emergency as this would negate tre requirement of first flying it to an airfield. The downsides of the RATO and parachute landings are that they require extra support equipment, such as a take-off and recovery vehicle, which increases the operational costs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002268_978-3-319-95603-9_8-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002268_978-3-319-95603-9_8-Figure5-1.png", + "caption": "Fig. 5 A schematic presentation of the configuration of CNT-modified ionic selective electrodes", + "texts": [ + " Ion-selective electrodes are commercially used to analyze many toxic ions, such as Hg2+, Cl\u2212, and F\u2212 in water samples [51]. The use of CNTs in ion-selective membrane electrodes can greatly enhance the sensor performance. The high electrical conductivity and surface area make CNTs potential components for use in field effect transistors (FETs) too [52]. In ionic selective electrodes, CNT can be used as an interface between the membrane and the inner electronic contact due to the fact that they do not undergo redox reactions and show good electrical conductivity (Fig. 5) [53]. On the other hand, a film of water can form at the membrane/SWCNT interface, which helps the ion-selective electrode response. In this light, many scientists have focused on the fabrication of CNT-modified ionic selective electrodes for the determination of water pollutants, especially toxic ions. As an example, Khani et al. (2010) described a novel potentiometric ion-selective electrode for the determination of Hg2+ as a toxic ion in water samples [54]. Modification of carbon paste electrodes (CPEs) with MWCNTs has been found to lead to a good potentiometric condition for the determination of Hg2+" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000243_imece2008-67408-Figure14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000243_imece2008-67408-Figure14-1.png", + "caption": "Fig. 14: Sensitivity analysis for local stiffness and global modes of the full trimmed body", + "texts": [], + "surrounding_texts": [ + "AND SENSITIVITY 6.1. Multi-body modeling A full vehicle multi-body model was build. The construction of the multi-body model followed a step-by-step procedure: \u2022 assembly of the front (Mc Pherson with flexible shock-absorber and front subframe) and rear (Twist Beam full flexible) suspensions; \u2022 assembly of the finite element trimmed body model , consisting in the implementation of: o front / rear suspensions; o steering system; o front / rear anti-roll bars; o powertrain; o front / rear wheels in Magic Formula representation; o full flexible trimmed body model of the vehicle in the two variants (Base and Enhanced). 6.2. Correlation of handling parameters with experimental results The correlation activity between experimental data and simulation results considering the full vehicle model in the two variants (Base and Enhanced) is then performed. A step-by-step correlation process is considered. Only main pseudo-static and dynamic maneuvers for the correlation are considered, i.e.: \u2022 Front and Rear suspension K&C characteristics; \u2022 CRC maneuvers at 40 m for Base and Enhanced variants; \u2022 ISO lane change maneuver at 90 km/h for Base and Enhanced variants. 8 aded From: http://proceedings.asmedigitalcollection.asme.org/ on 11/28/2017 Te In order to have confidence in the correlated models, a final verification of the other dynamic maneuvers, i.e.: \u2022 Step steer input at 100 km/h for Base and Enhanced variants; \u2022 Sinusoidal steering wheel angle input at 100 km/h for Base and Enhanced variants; has also performed. Copyright \u00a9 2008 by ASME Copyright \u00a9 2008 by ASME rms of Use: http://www.asme.org/about-asme/terms-of-use Download 6.3. Sensitivity analysis After getting a correlated model of the vehicle for the two variants, a sensitivity analysis of the main elastic and damping elements has been performed. Results of this analysis are not in the scope of this paper. In order to show the influence of the structural modifications, a modal methodology analysis is also applied. Through this methodology, effects of local (point mobilities) and global modes of the structures (front subframe, rear twist beam and full trimmed body) are considered. 9 ed From: http://proceedings.asmedigitalcollection.asme.org/ on 11/28/2017 T As a result of the sensitivity analysis and the experimental session, detailed indications have been provided to designers. Through them, optimized values for elastic, damping and structural components with respect to the global handling behavior of the car are defined. 7. CONCLUSION In order to asses the influence of the body flexibility on the handling behavior of a small size car, advanced analysis techniques combining a test setup with strain gauges attached to structural parts (front subframe, rear twist beam and body) and the application of the \u201cModal Contribution Methodology\u201d technique is applied. A full vehicle multi body correlated model was built. The model includes all vehicle subsystems: front and rear suspensions with rotating wheels and tire model in Magic Formula representation, steering line, powertrain mounting system and car body. Flexible components are also implemented, such as the trimmed body, the front subframe and the rear twist beam suspension. Sensitivity analysis based on elastic, damping and structural elements is performed to provide guidelines to designers. Copyright \u00a9 2008 by ASME Copyright \u00a9 2008 by ASME erms of Use: http://www.asme.org/about-asme/terms-of-use Downloa Results from the experimental measurements demonstrate that the methodology is suitable taking into account the influence of the body flexibility effect, which standard handling testing equipment do not allow. This work represents a preliminary study in the systematic approach of the body handling deformation. Further analyses will be necessary for having evidence of the advanced handling driver perception due to the body deformation and in defining a structured approach to correlate with objective measurements." + ] + }, + { + "image_filename": "designv11_92_0003692_s1068799818040104-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003692_s1068799818040104-Figure7-1.png", + "caption": "Fig. 7. Model of the bearing with deformable rings in NX Nastran.", + "texts": [ + " The rollers numbering is chosen such that the angular position of the first one is 1 =\u03d5 \u2212\u03c0 . First, let us analyze the bearing for the case of negative clearance (g = \u201340 \u00b5m). If the ring deformations are completely omitted, the obtained values of the contact forces turn out to be 2.5 times overestimated that can be seen in Fig. 6a. As all of the developed techniques are based on the same contact relation (6), in the case of rigid rings they all render identical results. The reference calculation was made with the aid of NX Nastran. To this end a 3D FEM model shown in Fig. 7 was constructed. The applied surface traction was distributed along the lower part of the inner ring raceway in accordance with (19). Kinematic boundary conditions of the outer ring are difficult to determine, so the latter is attached from both ends to long thin cylinders with a length of 235 mm. The opposite sides of cylinders are clamped. In order to vanish the cylinder contribution to rigidity of the assembly, their elasticity modulus is set = 0.01E E\u2032 . The FEM model of the bearing with rigid rings is a lot simpler", + "6 30.2 5.9 8.8 \u20130.8 5.4 \u03bbke \u2014pointwise BC 10.7 24.3 393.4 95.0 11.9 16.2 \u20133.3 12.3 \u03bbki \u2014pointwise BC 9.5 14.5 152.2 40.3 6.3 7.0 \u20130.4 5.4 The use of flexible foundation in the Dynamics R4 model was justified by the idea of imposing the kinematic boundary conditions as identical as possible to those used in the NX Nastran model, but other possibilities can be also considered. For instance, we also tried to replace the additional xxK stiffnesses by the pointwise constraints at points A and B (see Fig. 7) in the direction of the acting force, i.e. = = 0xA xBu u . Moreover, the same model was also applied to retrieve the numerical compliance coefficients for technique 3 and the results of the subsequent analysis are given in Fig. 9. RUSSIAN AERONAUTICS Vol. 61 No. 4 2018 577 A model of the radial roller bearing that accounts for structural deformations of its rings was presented in the paper. Several possibilities were developed for modeling ring flexibility. Each of the approaches proposed has its advantages and disadvantages and the final choice depends on the conditions the bearing operates at (negative or positive clearance, rotation speed value, etc) and the goal that the researcher attempts to achieve" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001804_0954406218770703-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001804_0954406218770703-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the ECTS in a shearer.", + "texts": [ + " The results indicate that the electromagnetic stiffness exerts a great influence on the bifurcation and chaotic motion. Nonlinear torsional dynamic model of the ECTS in a shearer To study the nonlinear torsional vibration characteristics of ECTS in a shearer, it is necessary to establish its practical nonlinear torsional vibration dynamic model. The model is a bridge between theoretical research and practical application, thus an accurate model should be built up to reflect the features and movement of the original system itself. The schematic diagram of the ECTS in a shearer is shown in Figure 1. The system components include PMSM, elastic torque shaft, coupling, reducer, cutting drum, and others. There are not only elastic parts but also inertial parts, so the ECTS in a shearer can be regarded as a complex multi-mass elastic system. To simplify the calculation without loss of generality and ignore the impact of the reducer gear gap, according to the mechanical principle, the ECTS in a shearer can be simplified into two inertia torsional vibration system, which consists of PMSM, cutting drum, and connection axis, as shown in Figure 2, where the moment of inertia of the gear portion is equivalent to the moment of inertia on the end cutting drum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000996_978-1-4020-6114-1_15-Figure15.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000996_978-1-4020-6114-1_15-Figure15.2-1.png", + "caption": "Fig. 15.2. Mobile hovercrafts and the error dynamics.", + "texts": [ + " Further, a novelty of this research that should not be overlooked is that the formation tracking in a 2-D setting studied in this Chapter includes hovercraft coordinating on a flat surface [7] or UAV flying at a constant altitude. Thus, the methodology proposed is easily extended and applied to UAV formation tracking in more general settings. A kinematics model of a hovercraft with two degrees of freedom is given by the following equations: cos , sin , x v y v (15.1) where the forward velocity v and the angular velocity are considered as inputs, (x, y) is the center of the rear axis of the vehicle, and is the angle between heading direction and x-axis as shown in Figure 15.2. For time-varying reference trajectory tracking, the reference trajectory must be selected to satisfy the nonholonomic constraint. The reference trajectory is hence generated using a virtual reference hovercraft [8] which moves according to the model: cos , sin , r r r r r r r rx v y v (15.2) where [xr yr r] is the reference posture obtained from the virtual vehicle. Following [8] the error coordinates are defined as (Figure 15.2): cos sin 0 sin cos 0 0 0 1 e r e e r e r x x x p y y y (15.3) It can be verified that in these coordinates the error dynamics become: cos sin . e e r e e e e r e e r x y v v p y x v (15.4) The aim of (single hovercraft) trajectory tracking is to find appropriate velocity control laws v and of the form: , , , , , , e e e e e e v v t x y t x y (15.5) such that the closed-loop trajectories of (15.4) and (15.5) are stable in some sense (e.g., uniform globally asymptotically stable). As discussed in Section 15" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000094_mawe.200800359-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000094_mawe.200800359-Figure7-1.png", + "caption": "Fig. 7. Schematic view of laser control system", + "texts": [ + " Again, algorithms need to be developed that provide the optimum temperature requirements for the components\u2019 thickness geometry and to determine the necessary energy levels and duration of exposure required at such levels. In addition, the algorithms will be devised for use in a computer program that will control laser energy through a thermal imaging feed back loop. A single laser beam can be divided and, thereby, provide a predetermined pattern of heat to the surface of an unformed blank and, subsequently, throughout the forming cycle. Furthermore, with computer control, variable heating patterns could be developed providing even greater control over material thickness distribution. In Fig. 7 a titanium disc is shown being heated by a single laser beam, divided into three separate beams, or heat sources. In this scenario, the unformed blank is preheated by two defocused beams to about 8500C, leaving the third computer controlled beam to raise the temperature to 9300C in those areas where full superplasticity is required. To achieve this, special mirrors and beam deflection devices are required. In Fig. 8, a 10kW laser, with an even energy distribution spot, is first directed at a mirror that allows 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002149_s40032-018-0480-4-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002149_s40032-018-0480-4-Figure1-1.png", + "caption": "Fig. 1 a Nomenclature of fir tree root [5], b contact pressure region in fir tree root", + "texts": [ + " Since the design applications of the contacting components are greatly influenced by their geometry, most often dovetail type of geometry composing of single tooth is adopted in rotors of compressors and fans. & A. Bharatish bharatisha@rvce.edu.in 1 Rashtreeya Vidyalaya College of Engineering, Bengaluru 560059, India In order to provide multiple point contacts and enhance the contact area leading to uniform stress distribution, fir tree type of geometry was also adopted composing of two or more lobes. Some of the geometrical specifications of fir tree rim are as shown in Fig. 1. The blade disc interface is mainly subjected to centrifugal, bending and thermal loads. This leads to the initiation of cracks in the fir tree rim region increasing the stress levels. Hence reduction of stress concentration requires utmost importance [3]. Therefore, some of the researchers have focused on analyzing fir tree joints in blade/disc assembly for achieving reduction in stress levels due to fretting fatigue. Papanikos et al. [1] reported that the flank angle, flank length, COF and skew angle hadsignificant effect on the stress distribution at blade/disc interface in aero engine compressor made of titanium (Ti-6Al-4V) alloy", + " It can be observed that the analysis error was found to be within the permissible limits for all the fretting responses and hence mesh convergence was ensured. Fretting parameters and their range were selected from the available literature [1, 4, 7] which were suitable for photoelastic stress freezing model. Taguchi\u2019s L27 orthogonal array was adapted to perform the simulations considering contact angle (15 , 17.5 , 20 ), speed (260, 360, 460 rpm) and COF (0.1, 0.3, 0.5). Coefficient of friction was considered between the fir tree root and inner face of the disc which is indicated by \u2018region of interest\u2019 in Fig. 1b. When the contact angle was below 15 , uniformity in the contact surfaces was not maintained whereas contact angle above 20 lead to geometrical distortions.When speed and COF was below 260 rpm and 0.1, respectively, significant reduction in the stress concentration was not observed. Also when speed and COF was above 0.5 and 460 rpm, respectively, the stresses exceeded the yielding limit leading to subsequent failure. Analysis of variance (ANOVA) was used to assess the significance of factors on fretting responses based on F-value obtained from F-distribution as shown in Table 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000939_sice.2008.4655035-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000939_sice.2008.4655035-Figure4-1.png", + "caption": "Fig. 4 A coordinate system of the 3D unicycle.", + "texts": [ + " Hence we get an orthogonal complement matrix Du2D to Cu2D for which vu2D = Du2D q\u0307u2D , Cu2D Du2D = 0, Du2D = \" I \u2212 \u201c C T u2D1 Cu2D1 \u201d \u22121 C T u2D1 Cu2D2 # . - 2231 - Hence we get the motion equation of the unicycle DT u2D Mu2D Du2D q\u0308 u2D = DT u2D \u201c hu2D \u2212 Mu2D D\u0307u2D q\u0307 u2D \u201d . (1) We focus on DT u2D hu2D in (3). This factor project hu2D to a tangent space of the unicycle. Thus we derived a motion equation of the unicycle which considered forces and torques to the unicycle by the operator using the Projection method. 3. MODELING OF A 3D UNICYCLE A unicycle consist rigid bodies which a wheel, cranks, pedals and a saddle in Fig.4. First of all, we derive motion equations of each rigid bodies. Finally, we get a motion equations to describe relations between each rigid bodies. 3.1 Modeling of sub-systems 3.1.1 A motion equation of a wheel A dynamic equation of a wheel, rolls without sliding on the horizontal plane, is Mwq\u0308w = hw, where is Mw = diag ` Ixzw + mwr2 w, Iyw + mwr2 w, Ixzw \u00b4 hw =2 4 gmwrwS\u03b8w + ` Iyw + mwr2 w \u00b4 \u03c9yw\u03c9zw \u2212 IxzwT\u03b8w\u03c92 zw \u2212mwr2 w\u03c9xw\u03c9zw \u03c9zw (\u2212Iyw\u03c9yw + IxzwT\u03b8w\u03c9zw) 3 5 . 3.1.2 The motion equation of the simple rigit bodes We derive motion equations of rigid bodes, saddle, cranks and pedals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure1-1.png", + "caption": "Fig. 1 Illustration of concentrated load applied to an annular plate [27]. (a) Concentrated load is applied to the rigid shaft; (b) equivalent axial load; (c) equivalent radial load", + "texts": [ + ", Kxy and Kyx) or radial stiffness (Kzz) terms cannot be represented. Thus, we use a thin plate to represent axial and radial stiffness instead of spring elements. 2.2 Model development Inspired by our previous work [27], we further develop axial and radial displacements on the basis of the axisymmetric annular plate model. A new stiffness model of bearings that is based on plate theory is proposed to replace commonly used spring elements. 2.2.1 Plate displacement An axisymmetric annular plate is illustrated in Fig. 1. The geometric parameters of the plate are as follows: Outer radius a, inner radius b, and plate thickness h. A concentrated load F, which includes axial load Fa and radial load Fr, is applied to the rigid shaft center O. The outer surface is fixed, and the inner surface is tied to a rigid shaft. 1) Axial deformation of the annular plate According to Liu et al. [28], the maximum axial displacement, wmax, of an axisymmetric annular plate under certain boundary conditions is a function of load (concentrated load F, uniform load Q, or uniform bending momentM), plate geometry (outer radius a, inner radius b, and thickness h), and material properties (Poisson\u2019s ratio \u03bc and modulus of elasticity E)", + " Under a uniform load, the maximum axial displacement is wmax \u00bc q E \u00bd f1\u00f0a,b,h, \u00de : (4) Under a concentrated load, the maximum axial displacement is wmax \u00bc F E \u00bd f2\u00f0a,b,h, \u00de : (5) Under a uniform bending moment, the maximum axial displacement is wmax \u00bc M E \u00bd f3\u00f0a,b,h, \u00de : (6) In the expressions above, fi (i = 1, 2, 3) are functions of geometric parameters and the Poisson\u2019s ratio of an annular plate. The corresponding function for different boundary conditions or loads is presented in Appendix. As shown in Fig. 1(a), the outer face of the annular plate is fixed, and the inner face is tied to a rigid shaft. Load is applied to the rigid shaft. The axial and radial displacements of the annular plate with applied load are very small; thus, the effect of axial displacement on radial stiffness can be disregarded and vice versa. Axial load Fa, which results in axial deflection wmax of the annular plate, can be equivalent to uniform load Qa and bending moment Ma because the inner face of the annular plate is tied to a rigid shaft (Fig. 1(b)). By adding up the axial deflections caused by Qa and Ma, we obtain maximum axial displacement wmax as follows: wmax \u00bc Qa E \u00bd f1\u00f0a,b,h, \u00de \u00fe M a E \u00bd f3\u00f0a,b,h, \u00de : (7) We can infer that Ma= BQa, where B is a constant that can be determined by a, b, h, and \u03bc, because the rotational effects of Qa and Ma on the inner face of the annular plate counteract each other. According to plate theory, we can infer that wmax \u00bc Q E 3b\u00f0 2 \u2013 1\u00de a4 \u00fe b4 \u2013 2a2b2 \u00fe 4a2b2ln a b ln b a 2h3\u00f0a2 \u2013 b2\u00de : (8) By substituting Q \u00bc Fa 2\u03c0b into Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.3-1.png", + "caption": "Fig. 11.3 Coordinate system and nomenclature. Reproduced from Ref. [2] with the permission of Guranpuri-Shuppan", + "texts": [ + " This book only discusses theoretical models. Table 11.1 summarizes theoretical models in terms of the shape of the contact pressure distribution, conditions of slip, bending deformation of the tread ring (belt), torsional deformation of the tire, translational lateral deformation of the tread ring, fore\u2013aft deformation of the whole contact patch, slip velocity/temperature dependency of the friction coefficient and transient temperature. The force and moment of a tire with slip angle a are expressed using the coordinate systems shown in Fig. 11.3. One coordinate system is fixed to the wheel while the other is fixed to the vehicle. It is necessary to understand which coordinate system is used in the definition of the force and moment of a tire. We denote the moving direction of the vehicle as the X\u2032-axis and the axis normal to X\u2032-axis as the Y\u2032-axis, while the direction of the wheel is the X-axis and the axis normal to the T ab le 11 .1 C ha ra ct er is tic s of m od el s di sc us se d in th is ch ap te r M od el pa ra m et er s So lid -t ir e m od el 11 ", + "45), the self-aligning torque in the adhesion region M0 z around the z-axis is given as M0 z \u00bc Cyl 2 hb lh 3 l 4 tan a\u00fe d\u00fe l2Cy 3r2Ky Fybl2h 4l2 lh l\u00f0 \u00de2: \u00f011:76\u00de The side force in the sliding region F00 y is given by F00 y \u00bc n\u00fe 1 n 2nFzld ln\u00fe 1 Z l lh l 2 n x1 l 2 n dx1 \u00bc n\u00fe 1 n 2nFzld ln\u00fe 1 l 2 n l lh\u00f0 \u00de 1 n\u00fe 1 l 2 n\u00fe 1 lh l 2 n\u00fe 1 ( )\" # : \u00f011:77\u00de The self-aligning torque in the sliding region M00 z around the z-axis is given by M00 z \u00bc n\u00fe 1 n 2nFzld ln\u00fe 1 Z l lh l 2 n x1 l 2 n x1 l 2 dx1 \u00bc n\u00fe 1 n 2nFzld ln\u00fe 1 1 2 l 2 n l 2 2 lh l 2 2 ( )\" 1 n\u00fe 2 l 2 n\u00fe 2 lh l 2 n\u00fe 2 ( )# : \u00f011:78\u00de The total side force Fy is obtained by adding the expressions in Eqs. (11.75) and (11.77): Fy \u00bc Cyl2hb 1 2 tan a d Cy \u00fe l2 3r2Ky Fy l 1 2 lh 3l \u00fe n\u00fe 1 n 2nFzld ln\u00fe 1 l 2 n l lh\u00f0 \u00de 1 n\u00fe 1 l 2 n\u00fe 1 lh l 2 n\u00fe 1 ( )\" # : \u00f011:79\u00de The total self-aligning torque Mz is obtained by adding the expressions in Eqs. (11.76) and (11.78): Mz \u00bc Cyl 2 hb lh 3 l 4 tan a\u00fe d\u00fe l2Cy 3r2Ky Fybl2h 4l2 lh l\u00f0 \u00de2 \u00fe n\u00fe 1 n 2nFzld ln\u00fe 1 1 2 l 2 n l 2 2 lh l 2 2 ( )\" 1 n\u00fe 2 l 2 n\u00fe 2 lh l 2 n\u00fe 2 ( )# : \u00f011:80\u00de Neglecting the rolling resistance, the cornering force FCF y in Fig. 11.3 is given by FCF y \u00bc Fy cos a: \u00f011:81\u00de The rolling resistance including the effect of the side force or the drag resistance Fdrag x is given by Fdrag x \u00bc Fy sin a\u00fe gFz cos a; \u00f011:82\u00de where \u03b7 is the rolling resistance coefficient. An iterative calculation is required to solve Eq. (11.79) because Eq. (11.79) contains Fy on both sides. The computational procedure is as follows. (i) Input l (contact length), b (contact width), Cy (lateral shear spring rate of the tread per unit area), Ky (lateral spring rate of the tire), Rmz (torsional spring rate of the tire), Fz (load), p (inflation pressure), n (parameter that determines the shape of the contact pressure distribution), EIz (in-plane flexural rigidity of the belt), \u03b7 (rolling resistance coefficient), r (tire radius), ls (static friction coefficient), ld0 (kinetic friction coefficient at zero sliding speed), aV (coefficient of the speed dependency of the kinetic friction coefficient), V (speed of the tire), a (slip angle) and Fy1 (initial guess of the side force)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002543_0954406218802604-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002543_0954406218802604-Figure3-1.png", + "caption": "Figure 3. FE model, meshing, constraints, and coordinate system.", + "texts": [ + " Coarse meshing is used in the distant area from the weld zone to save computational time. After each time step, the heat source moves towards the y direction by a distance equal to the product of the weld speed and time step. The time step and mesh size are selected in such a way that the distance by which the torch moves after each time step is greater than the mesh element length in the y direction. The complete mesh consists of 11,520 domains, 6368 boundaries, and 744 edge elements. The complete domain is divided into three regions (Figure 3). Fine meshing is used in the welding region (Region 2) and less fine meshing is adopted in the area of the temperature measurement points (Region 1). Coarse meshing is used in the distant area from the weld line (Region 3). The heat source is effective for 16.67, 8.34, and 5.56 s at welding speeds of 3, 6, and 9mm/s, respectively. After these periods, the workpiece continues to lose heat via convection and radiation until the temperature decreases close to room temperature. The commercial software COMSOL Multiphysics (v", + " This observation implies that thermal conductivity is the dominant variable and exerts more influence than specific heat. Decreasing the specific heat to a constant room temperature value does not increase the peak temperatures. The previous discussion focuses on the effect of material properties on transient temperature distribution at locations A and B, at the edge of the plate and away from the weld centerline. To investigate the effects on areas close to the weld, another location, C, is defined at a distance of 3mm from the weld line and adjacent to location A (Figure 3). The temperature distributions for different cases are plotted in Figure 8 to determine the change of trend in the material properties. At C, the maximum peak temperatures for Case 4 are observed at a weld speed of 3mm/s. At weld speeds of 6 and 9mm/s, the peak temperatures of Case 4 are underestimated by a small proportion, contrary to the peak temperatures at locations A and B. Figure 9 shows the stagnant temperature variation along the normalized longitudinal distance for the different cases at time\u00bc 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001198_1.2981120-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001198_1.2981120-Figure2-1.png", + "caption": "Figure 2. (A) Layout of the \u00b5-fluidic mold. (B) \u00b5-fluidic PDMS molding (C) System configuration.", + "texts": [ + " The chip comprises four cylindrical electrochemical chambers (2 mm, 1 mm, 0.5 mm, and 0.25 mm in diameter and 50 \u00b5m in height) with corresponding volumes of 157 nL, 39 nL, 9.8 nL, and 2.5 nL. Each cell contains three gold electrodes: working electrode (WE), counter electrode (CE) and reference electrode (RE) coated with Ag/AgCl layers (Figure 1). An Ag/AgCl open reference electrode was manufactured by a two stage electrodeposition process with a deposition rate of 0.57 \u00b5m/minute. The second part of the system includes \u00b5-fluidic polydimethylsiloxane (PDMS) molding (Figure 2B, see Figure 2A for the corresponding mold) and a measurement platform made of Delrin and a Perspex seal, both produced by CNC machining (see Figure 2C for the whole system configuration). Electrochemical characterization was experimented with an EmStat (PalmSens Inc.) potentiostat, measurement system and standard electrochemical assays. Ferrocyanide (10 mM), Ferricyanide (10 mM) and KCl (1 M) were mixed to yield an Fe(II)/Fe(III) solution. Concentrated solution of 8 mg/ml para-Aminophenol (pAP, Fw 145.6, Sigma) was diluted to final concentrations of 0.4, 0.04 and 0.004 mg/ml which were used for the cyclic voltammetry and chrono-amperometry (bias potential of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000294_14399776.2008.10785987-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000294_14399776.2008.10785987-Figure2-1.png", + "caption": "Fig. 2: Air pads type c : with sintered powders; d : with metallic mesh", + "texts": [ + " Interchangeable drilled inserts can be installed on both pad designs; for type b1, the feed pocket has a diameter d0 = 2 mm and depth \u03b4 = 1 mm, while for type b2 d0 = 4 mm and feed pocket depth \u03b4 can be varied. Fig. 1: Air pads with different supply systems, a: annular orifice; b1, b2: simple orifice with feed pocket Configurations of type b2 pad with d = 0.2, 0.3, 0.4 mm and \u03b4 = 10, 20, 30 \u03bcm were tested. With the same pad further investigations were conducted whose supply inserts were connected by a 30 \u03bcm depth and 0.8 mm wide circumferential groove. Figure 2 shows alternative supply systems with resistances consisting of different types of porous media: sintered metal powders (type c), see Belforte et al. (2007B) and fine-mesh woven wire cloth (type d), see Belforte et al. (2007A). In the first case, several cylindrical sintered bronze porous resistances with different diameters and particle sizes were housed in inserts installed in the pad. In the second case, the inlet resistance was produced using several types of commercial woven wire cloth with different porosity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002992_icems.2018.8549010-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002992_icems.2018.8549010-Figure4-1.png", + "caption": "Fig. 4. The conversion relationship between \u03b1-\u03b2 and g-h coordinates", + "texts": [ + " It is means that VTn1 must be opened before closing VTn2 when source Un is to be accessed into the system. VTn2 must be opened before closing VTn1 when source Un is to be removed from the system. The three basic vectors of multilevel technology and their duty cycle must be obtained firstly. To simplify the calculation of this process, the 60\u00ba coordinates system is applied instead of the \u03b1-\u03b2 coordinates system. The 60\u00ba coordinates system is assumed to be g-h coordinates system. The g axis coincides with \u03b1 axis. The h axis can be gotten by rotating g axis 60\u00ba counterclockwise. It is shown in Fig. 4. There is a conversion coefficients matrix to convert the vector in \u03b1-\u03b2 coordinates system into the vector in g-h coordinates system. It is shown in (9). Where the (Vrg,Vrh) is the coordinate of vector in g-h coordinates system, and (Vr\u03b1,Vr\u03b2) is the coordinate of vector in \u03b1-\u03b2 coordinates system. 11 3 20 3 rrg rrh vv vv (9) All coordinates of basic vector in g-h coordinates system become integer. For any reference vector Vref (Vrg,Vrh), the four closest basic vector can be obtained by rounding up and down the coordinates of the reference vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure7-1.png", + "caption": "Fig. 7 Minimum time of transition", + "texts": [ + " Although it is seemed to be much easier for a robot to play drums than to play trumpet or violin, there are some difficulties specific to the drum-playing robot. At first, since humanoid robots generally cannot move their own arms as quick as humans do, it takes too much time to move hands between drums during the drum performance. As a result of trial experiments by the use of our humanoid robots, it was found that moving a hand from the rightmost drum to the next takes 200[ms] and to the leftmost drum takes 280[ms] as in Fig. 7, although it is possible to beat a single drum at intervals of 70[ms]. Therefore, we decided to install only four minimum drums such as the snare drum, hi-hat, bass drum, and crash/ride cymbal on our robot as can be seen in Fig. 8. With respect to bass, the kick joint is added so as to beat it independently from the arm movement in the same manner as humans operate the bass drum with their foot. As shown in Fig. 9, the abovementioned four electric drums are connected to Trigger Module. A robot can send MIDI signal to Trigger Module through USBMIDI converter board" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003283_ciced.2018.8592365-Figure3-1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003283_ciced.2018.8592365-Figure3-1-1.png", + "caption": "Figure 3-1. Contact resistance heating device schematic", + "texts": [ + " In actual operation, the contact resistance of the busbar connector is small, so that the connector needs a certain heating power to reach a certain temperature, which requires a large current. It is difficult to simulate the actual scene of a large current passing through the joint under laboratory conditions. Therefore, we directly sandwich the electric heating tube in the middle of the joint, and replace the heating power when the large current passes through the contact resistance with the heating power of the electric heating tube. As shown in Figure 3-1. The contact resistance of the busbar connector is generally small. Assume that the contact resistance is 20u\u03a9. When the connector passed a current of 900 A, the heating power was 16.2 W. In the experiment, we directly used the heating tube to generate 16.2W of heating power. B. Determination of light heat and \u03b1 coefficient Illumination factors are also one of the factors that have a significant impact on joint temperature. In this experiment, we used high-power incandescent bulbs instead of outdoor sunlight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003742_apsipa.2018.8659673-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003742_apsipa.2018.8659673-Figure4-1.png", + "caption": "Fig. 4. Overall configuration of CPZ-ANF.", + "texts": [ + " (10) Since this system is interpreted as the digital filter for calculation of the instantaneous estimate of the gradient, throughout this paper HS(z) is referred to as the gradient filter. The magnitude responses of the gradient filter for \u03c1 = 0.9 is shown in Fig. 3. It is clear that the gradient filter has bandpass responses of which peak appears at the notch frequency \u03c90. In summary, the adaptive algorithm for update of a(n) is given by a(n+ 1) = a(n)\u2212 2\u00b5y(n)s(n). (11) The overall configuration of the CPZ-ANF is shown in Fig. 4. As stated above, the CPZ-ANF is one of the narrowband IIR filters. In general, the feedback loop of such filters has very large gain because their poles tend to be very close to unit circle. Therefore standard direct-form II structure suffers from high probability of overflow because the poles precede the zeros in series. Such overflow results in repeated generation of inaccurate output signals. In order to improve this problem, we analyze the magnitude responses of the feedback loop in the CPZ-ANFs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002890_j.matpr.2018.08.148-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002890_j.matpr.2018.08.148-Figure7-1.png", + "caption": "Fig. 7. Distribution of stress contour lines reduced based on H-M-H hypothesis in leaf no. 7.", + "texts": [ + " Due to the existing symmetry of the geometry and of the loading, the calculations were made only for a half of the model. Boundary conditions, enabling this type of the analysis of the structure were used. The loads acting between the leaves of the spring was considered by applying a proper contact connection. In order to verify a virtual model, comparison of the numerical and experimental tests was made. The authors compared obtained strains, displacements and the reaction forces. There was a broad convergence of the results (fig. 7). After adjusting the numerical model, the influence of the mounted rubber bumpers on the stress values was examined. Figure 8 presents the analysed object in the initial position and in the moment of maximal deflection of the rubber rolls. Obtained from the numerical simulation contour lines of the stress confirmed, that the cracks appeared on the object in the spots where the maximal stresses occurred. 26764 Sta\u0144co M., Iluk A., Dzia\u0142ak P./ Materials Today: Proceedings 5 (2018) 26760\u201326765 The crack of the shortest leaf occurred in the area near the mounting of a spring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003164_j.ifacol.2018.11.538-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003164_j.ifacol.2018.11.538-Figure2-1.png", + "caption": "Fig. 2. PCB design results and manufactured sample PCB", + "texts": [ + " In addition, an RC receiver, FM800, is connected to the microcontroller, and it is utilized during manual flight and to switch flight modes. Figure 1 shows a connection diagram of the components described above. The PCB design is completed by using EAGLE CAD software, and all PCBs are designed with four layers. The main PCB wheelbase is 130 mm and the dimensions of the sensor PCB is 35 \u00d7 35 mm. As was mentioned in Chapter 2.1, the ground plane is incorporated into the sensor PCB to improve GNSS signal reception. Figure 2 shows the PCB design results (left) and the manufactured sample PCBs (right); Fig. 3 shows the completed mini-drone, with all electronic parts soldered onto the PCBs and several parts installed, including the propeller, coreless motor, and RC receivers. The developed mini-drones are capable of autonomous flight in an outdoor environment; the detailed specifications are summarized in Table 2. IFAC SYROCO 2018 Budapest, Hungary, August 27-30, 2018 180 Dasol Lee et al. / IFAC PapersOnLine 51-22 (2018) 178\u2013183 Kalman filters (Kalman, 1960), extended Kalman filters, and unscented Kalman filters (Julier, 2004) are widely utilized for drone navigation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001614_access.2018.2811759-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001614_access.2018.2811759-Figure5-1.png", + "caption": "FIGURE 5. Shift region for pole placement.", + "texts": [ + " 4), where h \u2265 0 represents the prescribed degree of relative stability. This algorithm places closed-loop poles within the open sector using numerical iteration. VOLUME 6, 2018 15081 The best dynamic property of the control system will be obtained when the closed \u2013loop poles are placed near \u00b1\u03c0/2, but the largest region that can be considered is bounded by lines of \u00b1\u03c0/4. To better achieve the dynamic property of control system and increase the flexibility of the pole placements, a shifted sector method is proposed in Fig. 5, whereby the poles can be placed at any sector angle \u03b2 between 30\u25e6 and 90\u25e6. The design steps are as follows. Step 1) Given a linear system as in (53), assign h such that \u2212h is parallel with the imaginary axis, which would represent the line beyond which the closed-loop pole will be placed in the sector shown in Fig. 4. Assign the positive-definite matrix R, and solve the equation P0BR\u22121BTP0 \u2212 P0(A+ hEn)\u2212 (A+ hEn)TP0 = 0 (57) Solve the symmetric positive semi-definite matrix P0. The immediate closed-loop system matrix is A1 = A \u2212 BR\u22121BTP0, and hence all the poles are in the left-hand plane beyond the \u2212h vertical line. Set i = 1. Step 2) Assign an angle \u03b2 for the sector, and h1 = h in Fig. 5. If \u03b2 \u2208 (30\u25e6, 45\u25e6) go to Step3); otherwise \u03b2 \u2208 (45\u25e6, 90\u25e6) and proceed to Step7). Step 3) For assignment angle \u03b2, move the imaginary axis to the right: h2 = \u221a 3h1 tan\u03b2 \u2212 h1 (58) Obtain the new state matrix A1 = A1 \u2212 Eh2; Step 4). Solve the equation Q\u0302iBR \u22121BT Q\u0302i \u2212 Q\u0302i(A 3 1)\u2212 (A3 1) T Q\u0302i \u2212 0 = 0 (59) Obtain the symmetric positive semi-definite solution matrix Q\u0302i. Check if 0.5tr(BR\u22121BT Q\u0302i) = 0; if so, proceed to V, but otherwise continue. Step 5) Solve the equation P iBR\u22121BTP i \u2212 P(Ai)\u2212 ATi P \u2212 Q\u0302i = 0 (60) Obtain solution P i; the immediate closed-loop system matrix is therefore A\u0303 = Ai \u2212 \u03b3iBR\u22121BTP i", + " If so, proceed to V, but otherwise continue. Step 9) Solve the equation P iBR\u22121BTP i \u2212 PAi \u2212 ATi P \u2212 Q\u0302i = 0 (64) Obtain the solution P i; the immediate closed-loop system matrix is therefore \u02dcA = Ai \u2212 \u03b3iBR\u22121BTP i. Solve for the coefficient \u03b3i: \u03b3i = max{0.5, b2 + \u221a (b22 + a2c2 a2 } (65) where a2 = \u2212tr[(BR\u22121BTP i)2], b2 = tr(BR\u22121BTP i)Ai, and c2 = 0.5tr(BR\u22121BT Q\u0302i). Step 10) Set i = i+ 1 and proceed to Step8). Step 11) The algorithm is completed, all of the closed-loop system poles are in the specified region of Fig. 5, and the optimal feedback is K = BR\u22121BT (P0 + \u03b31P1 + \u00b7 \u00b7 \u00b7 + \u03b3jP j) (66) Finally, the optimal regulator can be given as u = \u2212Kx. NOTE 2: This pole placement technology, which is utilized to compute CAMF and the attitude manoeuvre controller, has an advantage over the traditional LQR algorithm; the feedbackmatrices can be obtained by setting the performance indexes without choosing a weight matrix Q, which is efficient and easy for high-dimension space state equations. The gain matrix designed by this algorithm is more convenient than the reference [15]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002164_042055-Figure1.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002164_042055-Figure1.1-1.png", + "caption": "Figure 1.1 Asymmetrical bearing-rotor system model diagram", + "texts": [], + "surrounding_texts": [ + "In this paper, a rotor system model with asymmetric support is established considering the nonlinear Hertz contact force of two kinds of rolling bearings. According to the principle of Lagrange equation, a set of dynamic differential equations of an asymmetrical rolling bearing-rotor system is established. 2 2 ( ) ( ) cos( t) ( ) ( ) sin( t) ( ) ( ) ( ) ( ) p p p p p p R p p L p p p p p p p R p p L p p R R R R p R p Rx R R R R p R p Ry R L L L L p L p Lx L L L L p L p m x c x k x x k x x m e m y c y k y y k y y m e m g m x c x k x x F m y c y k y y F m g m x c x k x x F m y c y k y y F Ly Lm g the bearing support force is: n x j,x b 1 1 1 n j,y b 1 1 1 cos( ) ( ) cos( ) sin( ) ( ) sin( ) Z Z Z j j j j j j j j Z Z Z y j j j j j j j j F F F K H F F F K H where i i o 2 (Z ) +j R t j R R Z is the angular position of the jth rolling element, cos sinj j jx y Gr is the contact deformation of the jth rolling element, 1 0 ( )= 0 0 j j j H \uff1e is the function of Heaviside, n is 3/2 when the rolling bearing is a ball bearing, and n is 10/9 when the rolling bearing is a cylindrical roller bearing. (1.1) (1.2) 3 1234567890\u2018\u2019\u201c\u201d \u03c9 rotating speed FLx, FLy the x and y direction contact force of cylindrical roller bearing" + ] + }, + { + "image_filename": "designv11_92_0003544_cac.2018.8623680-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003544_cac.2018.8623680-Figure2-1.png", + "caption": "Fig. 2. Relationship between Leader and Follower", + "texts": [ + " t ia is acceleration which controls the translation of UAV and n ia is heading angular velocity controlling UAV's deflection. The state variables and control variables are [ ]i i i ix y v\u03b8 and t n i ia a . B. Description of Positional Relationship between Leader and Follower The hypothetical conditions for UAV formation flight in this paper are as follows: 1) Ignoring the curvature of earth. 2) The ground coordinate system is considered as inertial coordinate system. The positional relationship between UAV i and the leader in formation is shown in Fig. 2. g g gx o y is ground coordinate system. b b b i i ix o y is body coordinate system of UAV i in which the mass center of UAV is the origin, the velocity direction of UAV is the positive direction of axis x , and the direction pointing to the right of UAV is the positive direction of axis y . Define 0 aP , iP and r jP as mass centers of the leader, UAV i and UAV j respectively, 0 aP and r jP are the points of attraction and repulsion of UAV i . a iP and r iP represent attraction bearing point and repulsion bearing point of The expression of each variable under the inertial system is as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003174_15440478.2018.1558149-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003174_15440478.2018.1558149-Figure1-1.png", + "caption": "Figure 1. Screw profile used in the experiments and sampling locations.", + "texts": [ + "16 kg) and a melting point around 160\u00b0C. Polypropylene grafted maleic anhydride (PP-g-MA) was used as compatibilizer. A ratio of PP-g-MA/fiber equal to one-tenth in weight was adopted to improve the affinity of the fibers with the matrix (Berzin, Vergnes, and Beaugrand 2014). A fiber content of 20% in weight was retained. The composites were prepared using a laboratory-scale twin-screw extruder (Clextral BC 21, Firminy, France). The extruder had a diameter of 25 mm and a length of 900 mm. The selected screw profile is shown in Figure 1. It was already used for other studies on twin screw extrusion of lignocellulosic fiber composites (Beaugrand and Berzin 2013; Berzin et al. 2017, Berzin, Vergnes, and Beaugrand 2014). The screw profile composed of one left-handed element (for polymer melting), screw conveying elements and blocks of kneading discs, a first one with a staggering angle of 90\u00b0, followed by a second one with an angle of \u221245\u00b0. The matrix (PP and PP-g-MA) was introduced in barrel 1 and melted before adding the fibers elements", + " The fibers were fed manually (in barrel 4), as done previously (Berzin, Vergnes, and Beaugrand 2014), to prevent possible irregular feeding, which usually occurs with long fibers when using either volumetric or gravimetric feeders. During all experiments, the barrel temperature was kept constant at 180\u00b0C, except for barrel P1 at 80\u00b0C, and barrel P2 at 120\u00b0C. The processing parameters were limited at a low screw speed (100 rpm) for a constant flow rate around 2 kg/h, in order to minimize fiber breakage. For fiber size monitoring along the screws, the feeding and screw rotation were abruptly stopped, the barrel was cooled down and extracted, then samples were taken at different locations. Figure 1 shows the axial distance of each sampling point along the extruder. A high-resolution Skyscan 1174 X-ray microtomograph with a closed X-ray micro-focus source was used for non-destructive three-dimensional (3D) image acquisitions. This scanner uses an X-ray source with adjustable voltage (20\u201350 kV), with a maximum power of 40 W, and a range of filters for versatile adaptation to different object densities. A sensitive 1.3 megapixel cooled CCD camera coupled to scintillator by lens with 1:6 zoom range, gives a spatial resolution of 6 up to 30 \u00b5m pixel size, with approximately 10 \u00b5m low-contract resolution", + " 2015): G K; \u03b1\u00f0 \u00de \u00bc 1 V \u03b3B\u03b1 K\u00f0 \u00de V K\u00f0 \u00de (7) where V(K) is the volume of the set K, \u03b2\u03b1 is the structuring element of size \u03b1 and \u03b3B\u03b1 K\u00f0 \u00de is the opening of K by \u03b2\u03b1 In the case of volume particle size analysis (for 3D images), if K is the set of fibers and \u03b2\u03b1 is an octahedron of size 1, the granulometry carries information regarding the distribution of the diameter (D) of the fibers defined as: G X; \u03bb\u00f0 \u00de \u00bc Pr D< 2 \u03bb\u00fe 1\u00f0 \u00def g (8) where D is the fiber diameter and 2 \u03bb\u00fe 1\u00f0 \u00deis the octahedron diameter. In this study, volume weighted diameter distributions were computed for different composite samples collected along the twin-screw extruder (Figure 1). Effects of processing on fiber shape deterioration Compounding flax fibers and PP matrix leads to uncontrolled fibre breakage mechanisms. In this regard, flax fiber-based composites content was assessed using X-ray tomography, which showed the dispersion of fibers, thereby giving indirect information on the type and degree of damage. Figure 4 shows a large variation in the size and shape of the raw fibers, when observed as delivered. Three main fibre elements can be defined: the elementary fibers consisting of a single long plant cell, the bundles consisting of parallel elementary fibers still glued together, and the fine particles having at most a 10 \u00b5m diameter", + " The shapes of the distributions of the fiber diameter during and after compounding are similar, showing a wide peak also around 40\u201380 \u00b5m for almost 30% of fiber population, and a range of diameters between 10 \u00b5m and 150 \u00b5m (Figure 7). The diameter distribution in Figure 7 shows that, on the one hand, fibers elements with D > 150 \u00b5m, arguably large bundles, disappear after compounding. More precisely, it seems that those bundles undergo the most damage at position P4, located on the second kneading block (see Figure 1), where only 2% of fiber elements kept a diameter around 140 \u00b5m. On the other hand, the frequency of fiber elements with the diameter of 10\u201360 \u00b5m increases progressively during the compounding process. As an example, only 3% of fiber elements with 30 \u00b5m in diameter was Figure 6. 3D reconstructed volume visualization of internal flax fiber elements distribution during compounding process (900 x 800 \u00d7 750 voxels, 1 voxel = 6.9 \u00b5m3). introduced as raw fibers at P1 position, for 27% of fibers at the extruder output (P5 position), which indicates that the final fibers are thinner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003432_oceans.2018.8604922-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003432_oceans.2018.8604922-Figure1-1.png", + "caption": "Fig. 1. Appearance and structure of floating ocean seismograph", + "texts": [ + " That is to say, self-adaptive fuzzy PID controller is more suitable for the seismograph in consideration of different control methods. And the designed controller needs to meet the following requirements: \u2022 In line with the depth control requirement of the floating marine seismograph, especially keep it in a certain depth. \u2022 Good performance with sufficient stability, short rise time and settling time, small overshoot and minimal steady-state error. II. FLOATING OCEAN SEISMOGRAPH STRUCTURE The floating ocean seismograph looks like an ocean bottom seismograph without pedestal, as shown in Fig. 1. 978-1-5386-4814-8/18/$31.00 \u00a92018 IEEE The variable buoyancy system is the main source for the floating ocean seismograph moving up and down. It acts as an actuator and plays a vital role in the process of depth control for the floating ocean seismograph. Because the maximum working depth of the floating ocean seismograph is about 1000m, the buoyancy adjustment capacity is 1.5L III. FORCES AND MODELING The static forces that the floating ocean seismograph receives are mainly buoyancy and gravity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure1-1.png", + "caption": "FIGURE 1 Gear-shaft interference assembly", + "texts": [ + " It allows connection of a gear with a shaft of comparable diameter where there is not enough material to design a spline connection at low cost. Due to this inherent motivation for the IF joint, the gear rim thickness is an important design consideration to avoid thin rim effects such as- effect of rim thickness on root stress, effect of rim thickness on crack propagation, and the effect of rim thickness on mesh stiffness. The IF gear is usually designed to have a back-up ratio i.e. the tooth whole depth to the rim thickness ratio greater than 1.2. A typical gear design has the back-up ratio value between 1.2 and 2. Figure 1 shows an interference-fit gear on a shaft. Table\u00a01 shows the geometry details of the gear and the shaft for the reference design discussed in detail and henceforth called design A. The gear-shaft joint in Figure 1 was analyzed by using Finite Element method. The gear-shaft joint was analyzed using a Frictional contact type joint in ANSYS Workbench software. As the gear-shaft joint and the effect on gear are the interest areas of the analysis, only the portion of the shaft close to the gear mounting was included in the finite element model. The shaft was fixed at the centerline at one end and there was no other force acting on the model. The static structural analysis was carried out for the gear-shaft interference set at nominal, minimum, and maximum interference" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003557_ichve.2018.8641899-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003557_ichve.2018.8641899-Figure1-1.png", + "caption": "Figure 1. Model of bottom phase transmission line", + "texts": [ + " Two types of composite insulators are arranged with lengths of 4 380 mm and 3 680 mm respectively. The supporting grading rings with external diameter 370 mm, pipe diameter 50 mm are equipped at the high voltage side and the conductor diameter is 26.8 mm with bundle distance 375 mm. For the reason that the non-uniform degree of electric field distribution at low phase is higher than the up one, the discharge will be more intense. In this paper, both the simulation and test research work for the lower phase conductor are mainly considered and simulation model is shown in Fig. 1. Referring to the original structure and parameters, the type of grading ring is half pipe with diameter 50 mm and equipped as the same direction. The applied voltage to the lower phase is 449.1 kV, which is the maximum phase to earth voltage value, and the voltage for the other two phases is half of the maximum phase to earth voltage value with the opposite direction. Electric field distribution result for grading ring is shown in Fig. 2. From Fig. 2, without any shielding part for two ending balls, the surface electric field intense for grading ring No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure25-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure25-1.png", + "caption": "Fig. 25 Schematic figures of two types of wire drive systems", + "texts": [ + " Similarly, an actuator pulls a wire cable which then pulls the robot frame. All the actuators are placed as close to the upper body as possible so as to reduce the moment of inertia of the leg about the knee and hip joint. The wire drive system makes possible such placement of actuators remotely from joints. Table 2 Types of wire drive systems Type 1 Type2 Mechanism One motor per joint One motor per cable Control Easy Complicated No. of motors One per joint Two per joint Tension High Low There are mainly two types of wire drive systems as in Fig. 25 and Table 2 [2, 14, 17, 39]. In many cases, to reduce the number of motors, the method using two wire cables and one motor, which is type 1, is often adopted. In this so-called \u201cpush-pull\u201d method, the load of wire cables is apt to increase since pretension of the cable is necessary. High cable tension causes several disadvantages: high friction between cables and guiding mechanical parts such as pulleys, low energy efficiency because of high friction, and short lifetime of cables. On the other hand, as for the \u201cpull-pull\u201d type (type 2), although two motors needed per each joint, the cable tension is controlled by the motors and the cable load can be lowered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000413_20080706-5-kr-1001.00300-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000413_20080706-5-kr-1001.00300-Figure2-1.png", + "caption": "Fig. 2. Schematic free-body diagram of four-rotor minihelicopter.", + "texts": [ + " In section 2, the model and basic properties in discrete-time are given. In section 3, the main results are presented; namely, the local exponential observer and the linearizing control law are designed, and a kind of separation principle is derived. In section 4, some real-time experiments are presented. Finally in section 5 some conclusions are given. 978-3-902661-00-5/08/$20.00 \u00a9 2008 IFAC 1753 10.3182/20080706-5-KR-1001.3589 2.1 Continuous-time model. A schematic free-body diagram of VTOL is given in fig. 2. The dynamic model is quite standard, the reader can see e. g. Castillo et al. [2007], Madani and Benallegue [2006], Samir et al. [2004], Tayebi and McGilvray [2004]. Is this paper the yaw angle \u03c8 is assumed to be zero and the model includes the remote-control parameters. Then the simplified model is given by x\u0308 = u1 m sin \u03b8 cos\u03c6 \u03b8\u0308 = \u2212\u03b11 Iy \u03b8\u0307 \u2212 \u03b12 Iy \u03b8 + \u03b13 Iy u3 y\u0308 = \u2212u1 m sin\u03c6 (1) \u03c6\u0308 = \u2212\u03b21 Ix \u03c6\u0307\u2212 \u03b22 Ix \u03c6+ \u03b23 Ix u2 z\u0308 = u1 m cos \u03b8 cos \u03c6\u2212 g where ui, i = 1, 2, 3 are the control inputs,m, Ix and Iy are the mass and moments of inertia of the VTOL" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002887_ihmsc.2018.00016-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002887_ihmsc.2018.00016-Figure3-1.png", + "caption": "Figure 3. Schematic diagram of the balance of force in hover mode.", + "texts": [ + " The resultant force of thrust, lift and drag is supposed to be vertical to offset part of gravity and thus achieving short takeoff and landing [6], as shown in Figure 2. The shafts of main propellers are parallel with the centerline of the fuselage. In order to get a vertical force for take-off and landing, a certain angle between the body and the ground shall be maintained. As proved by previous research, it is feasible to maintain this angle by installing a propeller vertically at the nose of to the aircraft. This arrangement enables the angle control and moment balance before take-off, as can be seen in Figure 3. III. MATHEMATICAL MODEL OF THE V/STOL This V/STOL aircraft involves a vertical tension for the balance of force, and the impact of the propeller deflected slipstream shall be considered. As the result, the traditional longitudinal dynamic equation in body axes shall be rewritten as follows: Force equations: (1) (2) Where , are the propeller deflected slipstream forces. Moment equation: (3) Where is the angular velocities, is the moment of propeller deflected slipstream. Dynamic equations for identification: (4) (5) (6) Considering the relationship among the propeller deflected slipstream, airspeed, propeller tension and control surfaces, the model equations can be written as follows: Force and moment equations: (7) (8) (9) Linear nondimensional aerodynamic force and moment equations are given below: (10) (11) (12) Where is based on the parameters for level flight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003394_6.2019-1747-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003394_6.2019-1747-Figure4-1.png", + "caption": "Figure 4: Examples of displacements implemented in ABAQUS where the symmetric angle is \u00b160 deg.", + "texts": [ + " Four node shell (S4R) elements are meshed on the hinge shell using a 1 mm mesh. The asymmetric configurations are implements as displacement and rotation boundary conditions 6 OF 14 AMERICAN INSTITUTE OF AERONAUTICS AND ASTRONAUTICS in static/general steps. Each range of asymmetric configurations is explored as a separate step enforced on an initially symmetric configuration. An asymmetric data set is generated for each primary fold angle, \u03b81, at increments of 5 degrees, resulting in 16 equal sense and 10 opposite sense data sets for each material. Figure 4 shows example profiles for the equal sense and opposite sense cases and with non-symmetric deviations, with a no added deformation scaling. Designing the displacement and rotation boundary conditions such that the simulations converge without error is not trivial and not easily automated. The approach here is to fix the inbound hinge frame to zero displacements and to apply displacements and necessary degrees of freedom to the outbound frame. Then the reaction forces, reaction moments, displacement, and rotational displacements are reported for the reference points representative of the hinge reference frames" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000213_s11665-008-9314-5-Figure13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000213_s11665-008-9314-5-Figure13-1.png", + "caption": "Fig. 13 (a) Example of a fracture surface of Alloy 3. (b) Small area of crack propagation is observed", + "texts": [ + "0 that calculates the Weibull s statistical distribution. The specimens 2 and 5 from Alloy 2, the tests were interrupted because both lives were superior to 107 cycles (runout). Table 6 presents the life results for B10 and B90 probabilities. Obs: *Specimens 2 and 5 from Alloy 2 the testing was interrupted by reaching the projected run-out (107 cycles) Dotted line (\u00c6\u00c6\u00c6) means that the samples were lost due to overloading or other testing problem The fractographic analysis of the surface fracture, Fig. 13, showed two fronts of slow crack propagation, nucleated at the root of the thread, which presents the highest local stress, and the growth took place in semi-elliptical form. From Table 5, it is concluded that Alloy 2 presented the best results in fatigue, probably due to its chemical composition that provided the best association between ductility and mechanical strength, since the threads are produced by a forming process. This work showed that it is possible to develop low alloy steels and a cold work process (without the use of heat treatments) for U-bolts used in the leaf suspension springs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003710_978-3-030-13317-7_8-Figure8.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003710_978-3-030-13317-7_8-Figure8.5-1.png", + "caption": "Fig. 8.5 Bering element with artificial damage in the outer race", + "texts": [ + " The 4:1 ratio gearbox consists of two gears, the driver gear has 18 teeth and the driven gear has 72 teeth; thus, the wear has been artificially induced uniformly by a gear factory in all teeth of three similar driven gears. From Fig. 8.4a\u2013d are shown the set worn gears analyzed in this proposal: healthy and 25, 50 and 75% of uniform wear, respectively. In regard with the damaged bearing, the IM uses a bearing model 6205-2ZNR that is located in rotor on the output shaft side; thus, a similar bearing element has been also artificially damaged by means of drilling a through-hole on its outer race with a tungsten drill bit of 1.191 mm diameter to produce the bearing with damage. In Fig. 8.5 is shown a picture of the damaged bearing used to perform the proposed analysis. This faulty condition is analyzed due to most of the rotating machinery involves the use of bearing elements, and their sudden malfunctioning tend to introduce nonlinear effects in the whole elements that are indirectly linked to bearings with damage. During the experimentation, two different operating frequencies are set in the VDF to drive the IM, that is, 15 and 50 Hz. Such operating frequencies produce an averaging output rotating speed of 889 rpm and 2984 rpm, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003119_s1063785018110196-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003119_s1063785018110196-Figure1-1.png", + "caption": "Fig. 1. (a) Optical scheme of the laser cladding device and (b) multiple-pass scheme. (1) Laser, (2, 5) rotary mirrors; (3, 4, 8, 12) conical mirrors, (6) the tube feeding the deposited material, (7) lens, (9) cladding material, (10) substrate, (11) cylindrical mirror, and (13, 14) circular laser beams.", + "texts": [ + " The initial circular laser beam is divided into several annular beams with a regulated distribution of laser power over them. The circular beams are converted into conical ones, which are focused separately on the substrate surface and on the material being cladded (powder jet, wire, slurry jet) for heating. The foci of conical beams are located along the optical axis, along which the cladding material is fed. A number of devices have been developed to implement the method. One device works as follows (Fig. 1a). After having passed through the regulated beam expander, laser beam 1 is converted into two circular beams. One of the beams is converted into a conical one by lens 7 and focuses on surface 10 into an irradiation spot. The applied material is fed through tube 6 in the form of stream 9 of powder or wire. Another conical beam is focused by means of conical mirrors 8 in given f low areas 9 to heat them. A special optical system was developed for the formation of a series of circular laser beams with regulated distribution of laser energy over the circular beams", + " At the initial stage of heating the particles, the evaporation and thermal conductivity losses are small and the temperature dynamics can be written in the form (6) According to calculations, the required laser power for heating to 1500 K does not exceed \u223c200\u2013300 W at a duration of the stay of the particles in the heating region of \u223c10\u20132 s. At a low density of particles in the powder stream, the efficiency of laser radiation absorption drops sharply at a single passage of radiation through the stream. A multipass optical scheme has been proposed [6] using cylindrical 11 and conical 12 mirrors to maximize the use of laser energy (Fig. 1b). First, the heating of the stream increases. Second, the absorption of laser radiation increases up to the full absorption when the coherent part of the laser radiation is returned to the laser resonator by mirror 12. Laser wire cladding is a process that has many benefits over traditional clad processes. Usually, wire feeding is performed both from the side and coaxially when using a conical beam [4]. This approach uses focus spots of millimeter size and laser power at the level of a few kilowatts to heat both the substrate and the wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000313_j.chaos.2008.07.018-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000313_j.chaos.2008.07.018-Figure1-1.png", + "caption": "Fig. 1. (a) Preloaded bearings mounted on shaft and (b) the corresponding mathematical model.", + "texts": [ + " As the preload is to be kept as low as possible the risk of entering a situation where not all rollers are in contact with a raceway increases. The corresponding bifurcation, which then can occur, is critical for the wear of the bearing and shall hence be included in the mathematical model. In this paper, a model of spherical roller thrust bearings previously introduced by Karlberg and Aidanp\u00e4\u00e4 [14] is used to study transitions and a saddle node bifurcation. The objective for this paper is to develop a tool for choosing a suitable preload, and to show that a saddle node bifurcation is present at certain sets of parameters. Fig. 1a shows a physical model of a rotating shaft supported by roller thrust bearings (RTB). A constant preloading Fp is applied in axial direction with the aim to keep all rollers in contact with the raceway (full contact). For simplicity, the rollers are considered flexible while the other parts are assumed stiff. In order to reduce the state space only translations in i and j directions are considered. When the radial displacement of the shaft exceeds the radial compression of the rollers created by the preloading some of the rollers will loose contact. When this occurs the bearing stiffness will rapidly decrease since some of the springs (rollers) becomes inactive. Hence, in this model it is assumed that the resulting stiffness in radial direction will decrease by fifty percent when this occurs. Fig. 1b shows to the Fig. 1a corresponding mathematical model derived from the assumptions above. A rigid shaft with mass m, spin speed X, and eccentricity c, is connected to the housing by the RTB modelled as springs with stiffness kr, and linear viscous dashpots with damping constant c. The radial position vector r has the components x and y in the direction i and j, and the length of this vector is denoted r. The model also includes gravity which is denoted g while the phase angle is denoted u. Karlberg and Aidanp\u00e4\u00e4 [14] showed that by introducing the nondimensional quantities x\u0302 \u00bc ~x c ; y\u0302 \u00bc ~y c ; t\u0302 \u00bc t ffiffiffiffiffiffiffiffiffiffiffi kr=m q ; bX \u00bc Xffiffiffiffiffiffiffiffiffiffiffi kr=m p ; f \u00bc c 2 ffiffiffiffiffiffiffiffi krm p ; r\u0302 \u00bc r c and bF pr \u00bc Fpr krc \u00f01\u00de the governing nondimensional equations of motion to the model in Fig. 1 (without gravity) become x\u030200 \u00fe 2fx\u03020 a\u0302x\u0302 \u00bc bX2 cos\u00f0bX t\u0302\u00de \u00f02\u00de y\u030200 \u00fe 2fy\u03020 a\u0302y\u0302 \u00bc bX2 sin\u00f0bX t\u0302\u00de \u00f03\u00de where a\u0302 \u00bc 2 r\u0302 < bF pr \u00f01\u00fe bF pr=r\u0302\u00de r\u0302 P bF pr ( \u00f04\u00de If r\u0302 < bF pr all rollers are in contact with the raceway (full contact) which is modelled as contact with two preloaded springs and hence a\u0302 \u00bc 2 (the preload of the two springs cancel each other). If on the other hand r\u0302 P bF pr some of the rollers will loose contact with the raceway (partial contact) which is modelled as contact with one preloaded spring and hence a\u0302 \u00bc 1 bF pr=r\u0302. A prime (0) represents derivatives with respect to the nondimensional time t\u0302 i.e. d dt\u0302 . By substituting the equations x\u0302 \u00bc r\u0302 cos\u00f0bXt\u0302 u\u00de \u00f05\u00de y\u0302 \u00bc r\u0302 sin\u00f0bXt\u0302 u\u00de \u00f06\u00de into Eqs. (2)\u2013(4) which results in the new equations of motion r\u030200 \u00fe 2fr\u03020 \u00f0a\u0302\u00fe \u00f0X u0\u00de2\u00der\u0302 \u00bc bX2 cos\u00f0u\u00de \u00f07\u00de r\u0302u00 2\u00f0r\u03020 \u00fe fr\u0302\u00de\u00f0bX u0\u00de \u00bc bX2 sin\u00f0u\u00de \u00f08\u00de equations of motion can be expressed as functions of phase angle u (defined in Fig. 1) and radial displacement r. By the substitutions vr \u00bc r\u03020 and vu = u0 Eqs. (7) and (8) can be written in state form as r\u03020 \u00bc vr \u00f09\u00de v0r \u00bc bX2 cos\u00f0u\u00de 2fvr \u00fe \u00bda\u0302\u00fe \u00f0bX vu\u00de2 r\u0302 \u00f010\u00de u0 \u00bc vu \u00f011\u00de v0u \u00bc \u00bd bX2 sin\u00f0u\u00de \u00fe 2\u00f0vr \u00fe fr\u0302\u00de\u00f0bX vu\u00de =r\u0302 \u00f012\u00de In vector notation these equations can be written as r\u03020u \u00bc F\u00f0r\u0302u\u00de \u00f013\u00de where r\u0302u \u00bc r\u0302 vr u vu T \u00f014\u00de The equations of motion expressed in r\u0302 and / (see Eqs. (7) and (8)) are autonomous which hence gives an opportunity to analyse the existence of stationary points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure6-1.png", + "caption": "FIG. 6: CAD MODEL, AND A PLANAR BUILD STRATEGY IS EMPLOYED TO MANUFACTURE A STOCK MODEL, WITH THE TWO BOOLEAN CUT OPERATIONS FOR THE HOLES SUPPRESSED.", + "texts": [ + " This will reduce interference conditions. Model preparation needs to be done to facilitate the build process. Manual spatial decomposition must be performed, where build features, build planes, and operation sets are defined by the planner. Features that required support material for a given build orientation should be suppressed (for a solid model), or construction geometry added as appropriate. The \u2018flanged bullet\u2019 in Fig. 2 (case (i)) has the holes suppressed, and a planar build solution illustrated in Fig. 6. 4 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use This generated stock model is utilized for the finish machining. It is assumed to be fixtured using the square flange, and an outer diameter (OD) turning tool path is using to remove 1 mm of material. A \u00be inch (19.05 mm) tool is used for drilling, and 24 mm boring tool are used to produce the holes. The virtual model finished model of the component is illustrated in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001134_2008-01-2623-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001134_2008-01-2623-Figure5-1.png", + "caption": "Figure 5: Mode at 58.7 Hz (to be excited at ~900rpm - 4th order or harmonics of 2nd order, ~1800rpm - 2nd order) \u2013 Strain Energy Density Contour", + "texts": [], + "surrounding_texts": [ + "A modal analysis is performed first in order to obtain the critical modes of the bracket. Range for modal analysis (0-900Hz) is chosen to cover twice the excitation frequency, which is 6th order engine load at 4500 rpm corresponding to 450 Hz. The system has 113 modes between 0-900 Hz. Critical modes of the bracket are at 18.6Hz, 29.4Hz, 58.7Hz, 352.8Hz, 402.5Hz, 410.1Hz, 566.3Hz and 803.4Hz. Strain energy density contours for the critical modes of the bracket are shown in Figures 3 \u2013 10. 276 SAE Int. J. Commer. Veh. | Volume 1 | Issue 1" + ] + }, + { + "image_filename": "designv11_92_0003819_b978-0-12-812939-5.00002-1-Figure2.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003819_b978-0-12-812939-5.00002-1-Figure2.1-1.png", + "caption": "FIG. 2.1 Planar two-linkmanipulator with two actuated jointsO1,O2 in gravitational field andwith jointO1 affixed to lab inertial frame. Joints angles, \u03b81, \u03b82, links dimensions L1, L2, center ofmasses, CM1, CM2, proximal distances to center of masses LCM1, LCM2, links massesm1,m2, moments of inertia about links\u2019 center of masses I1, I2, and end-effector OEE are depicted.", + "texts": [ + " Still further, readers may also opt to completely skip this chapter and without much consequence and overdue, focus their attention on other material presented in this book. When needed, readermay clearly return to Chapter 2 for clarification, or look into one of themany useful references on these topics. 2.2 Kinematics Probably, the simplest example of dynamic system that would suffice to illustratemajority of the concepts we intend to cover in this brief crash course onmanipulator (and locomotor) kinematics and dynamics is a planar two-link manipulator in gravitational field depicted in Fig. 2.1. This system has two actuated joints: joint 1 in between ground and link 1, and joint 2 in between link 1 and link 2. Joint 1 is attached to stationary point in the lab inertial frame. We refer to the nonactuated end of manipulator as end effector. One may draw a parallel between these two-link manipulator and human arm constrained to move in sagittal plane. Joints O1 and O2 relate to shoulder and elbow, and end effector relates to wrist and fist. 2.2.1 Forward Kinematics Forward kinematics refers to process of obtaining position and velocity of end effector, given the known joint angles and angular velocities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003675_022092-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003675_022092-Figure10-1.png", + "caption": "Figure 10(a), the velocity cloud map of the oil film, shows that the speed of oil film in the intersection area between the fluid and the solid is significantly greater than any other areas, including the free surface. This is because movement of the lubricating oil caused by the movement of the inner ring of the bearing and the rolling elements; the oil film in the contact area is thinner, hence better entrainment effect and larger flow rate. As oil film gets thinner from the inlet to the outlet of the contact area, the lubricating oil moves toward this area; as it gets thinker after bypassing the contact area, most of the lubricating oil bypasses the contact area. Analysis of the less loaded part of the rolling element is shown in Figure 10(b), which shows that the velocity of the oil film at the inlet is greater than that at the outlet. Analysis of the less loaded part of the rolling element is shown in Figure 10(c), which also shows that the oil film velocity at the inlet is larger than that at the outlet. That is because as the oil film in the contact area gets thinner, more lubricating oil accumulates at the inlet and less at the outlet. Figure 10 (d) is a radial sectional view of the oil film.", + "texts": [], + "surrounding_texts": [ + "(a) Overall images of oil film flow rate (b) Contact load less scroll body oil film flow ges CISAT 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1168 (2019) 022092 IOP Publishing doi:10.1088/1742-6596/1168/2/022092 5.Conclusion This paper takes the bearing NJ2232 at the output end of the rocker arm of thin coal seam shearer as the research object to establish the bearing lubrication model. Based on contact theory and elastic fluid dynamics lubrication theory, this paper analyzes the lubrication effect of low-speed heavy-duty bearings. The conclusions are as follows: (1) The fluid-solid coupling model of the bearing is established. Selected parameters of the bearing and bearing lubricant and contact conditions are applied to the simulation model, thus verifying the correctness of the fluid-solid coupling model. (2) The lubrication performance of the contact interface of low-speed heavy-duty bearings is analyzed. Results show that the thickness of the oil film features rectangular distribution in the cross section; the rectangular area of the oil film thickness between the rolling element and the inner ring is smaller than that between the outer ring and the outer ring. The oil film pressure at the inlet of the contact zone increases, so the pressure at the inner rolling contact is greater than that at the outer raceway. The velocity of the oil film at the intersection part between the fluid and the solid is greater CISAT 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1168 (2019) 022092 IOP Publishing doi:10.1088/1742-6596/1168/2/022092 than that of the free surface and is significantly greater than any other areas. In different working conditions, as the rotational speed increases, the thickness and pressure of oil film increase; as the load increases, oil film thickness decreases while oil film pressure increases." + ] + }, + { + "image_filename": "designv11_92_0001017_cira.2009.5423180-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001017_cira.2009.5423180-Figure1-1.png", + "caption": "Fig. 1. Mobile robot navigation in global coordinate.", + "texts": [ + " But since the encoder measurement is based on the wheel shaft, it leads inevitably to unbounded accumulation of errors if wheel slippage occurs. Specifically, orientation errors will cause large lateral position errors, which will increase proportionally with the distance traveled by the robot. As described in the previous paragraph, one can estimate the position and heading angles of the mobile robot using the feedback information from the two encoders on the left and right wheels of mobile robot shown in figure 1. The distance and heading increment can be obtained as follows: , , 2 en en R k L ken k \u0394d \u0394d \u0394d (1) , , en en R k L k k k \u0394d \u0394d \u0394 B \u03c8 (2) Then the position (X, Y, \u03c8) can be estimated as 1 , en k k x kX X d , (3) 1 , en k k y kY Y \u0394d (4) 1k k k\u0394\u03c8 \u03c8 \u03c8 (5) , where , cosen en x k k k\u0394d \u0394d \u03c8 , sinen en y k k k\u0394d \u0394d \u03c8 The analog light transmitter makes it possible to convert analog light values. The light value can be transmitted in parallel on 8 channels (sensors from Inex company shown in figure 2), which are independently programmable, and can thereby be compared to 8 different threshold values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002647_jitr.2019010110-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002647_jitr.2019010110-Figure4-1.png", + "caption": "Figure 4. The Expression of Ship Motion Standard Symbols in Inertial Coordinate System and BODY System", + "texts": [], + "surrounding_texts": [ + "The\ufeff BODY\ufeff coordinate\ufeff system\ufeff is\ufeff an\ufeff accompanying\ufeff coordinate\ufeff system\ufeff that\ufeff accompanies\ufeff the\ufeff movement\ufeffof\ufeffthe\ufeffship.\ufeffThe\ufeffcenter\ufeffcoordinates\ufeffare\ufeffgenerally\ufeffdefined\ufeffat\ufeffthe\ufeffcenter\ufeffof\ufeffthe\ufeffship\u2019s\ufeffwaterline.\ufeff The\ufeffsymbolic\ufeffrepresentation\ufeffof\ufeffthe\ufeffBODY\ufeffcoordinate\ufeffsystem\ufeffis\ufeff b x y z\nb b b{ } = ( ), , .\ufeffThe\ufeffx-axis\ufeffof\ufeffthe\ufeff coordinate\ufeffsystem\ufeffrepresents\ufeffthe\ufeffrolling\ufeffaxis,\ufeffwhich\ufeffis\ufeffthe\ufeffdirection\ufefffrom\ufeffthe\ufeffstern\ufeffto\ufeffthe\ufeffstem.\ufeffThe\ufeff y-axis\ufeffis\ufeffthe\ufeffpitching\ufeffaxis,\ufeffwhich\ufeffis\ufeffthe\ufeffdirection\ufefffrom\ufeffthe\ufeffport\ufeffside\ufeffof\ufeffthe\ufeffship\ufeffto\ufeffthe\ufeffstarboard\ufeffside.\ufeff The\ufeffz-axis\ufeffrepresents\ufeffthe\ufeffyawing\ufeffaxis,\ufeffwhich\ufeffis\ufeffthe\ufeffdirection\ufefffrom\ufeffthe\ufefftop\ufeffof\ufeffthe\ufeffship\ufeffto\ufeffthe\ufeffbottom\ufeff of\ufeffthe\ufeffship.\ufeffThe\ufeffsymbols\ufeffgiven\ufeffin\ufeffTable\ufeff1\ufeffindicate\ufeffthe\ufeffposition\ufeffand\ufeffdirection,\ufeffforce\ufeffand\ufefftorque,\ufeffand\ufeff linear\ufeffand\ufeffangular\ufeffvelocities\ufeffof\ufeffthe\ufeffship\u2019s\ufeffmotion\ufeffin\ufeff6\ufeffdegrees\ufeffof\ufefffreedom.\nIn\ufeffthe\ufeffstudy\ufeffof\ufeffthe\ufeffdifferent\ufefftypes\ufeffof\ufeffship\ufeffmotion\ufeffcontrol\ufeffalgorithm,\ufeffthe\ufeffmathematical\ufeffmodel\ufeffof\ufeff ship\ufeffmotion\ufeffmodeling\ufeffmethods\ufeffare\ufeffalso\ufeffdifferent.\ufeffPeople\ufeffusually\ufeffchoose\ufeffsingle\ufeffdegree\ufeffof\ufefffreedom\ufeff mathematical\ufeffmodel\ufeffwhen\ufeffstudying\ufeffthe\ufeffship\ufeffheading\ufeffcontrol\ufeffalgorithm.\ufeffThis\ufeffmodel\ufefftakes\ufeffthe\ufeffsteering\ufeff angle\ufefffor\ufeffa\ufeffcontrol\ufeffinput\ufeffand\ufeffthe\ufeffheading\ufeffangle\ufefffor\ufeffthe\ufeffoutput-feedback.\ufeffPeople\ufeffusually\ufeffchoose\ufeffthree\ufeff degrees\ufeffof\ufefffreedom\ufeffmathematical\ufeffmodel\ufeffwhen\ufeffstudying\ufeffthe\ufeffship\ufefftrack\ufeffcontrol\ufeffalgorithm.\ufeffThis\ufeffmodel\ufeff", + "takes\ufeffthe\ufeffcommand\ufeffsteering\ufeffangle\ufeffand\ufeffthe\ufeffcommand\ufeffthrust\ufefflevel\ufefffor\ufeffa\ufeffcontrol\ufeffinput\ufeffand\ufeffthe\ufeffsurging,\ufeff swaying\ufeffand\ufeffyawing\ufeffof\ufeffthe\ufeffship\ufefffor\ufeffthe\ufeffoutput-feedback.\nAccording\ufeffto\ufeffIEC62065\ufeffstandards,\ufeffthe\ufeffship\ufeffmotion\ufeffmathematical\ufeffmodel\ufeffdesigned\ufeffin\ufeffthis\ufeffpaper\ufeff mainly\ufeffincludes\ufefffive\ufeffparts:\ufeffthe\ufeffresponse\ufeffmodel\ufeffof\ufeffpropulsion\ufeffunit,\ufeffthe\ufeffresponse\ufeffmodel\ufeffof\ufeffsteering\ufeff gear,\ufeff the\ufeff response\ufeffmodel\ufeffof\ufeff surging,\ufeff the\ufeff response\ufeffmodel\ufeffof\ufeff swaying\ufeffand\ufeff the\ufeff response\ufeffmodel\ufeffof\ufeff yawing.\ufeffThe\ufeffcomposition\ufeffof\ufeffthe\ufeffmodel\ufeffis\ufeffshown\ufeffin\ufeffFigure\ufeff5.\ufeffThe\ufeffestablishment\ufeffof\ufeffship\ufeffmotion\ufeffmodel\ufeff is\ufeffshown\ufeffas\ufefffollows:\nResponse Model of Propulsion Unit The\ufeffpropulsion\ufeffunit\ufeffcontrolled\ufeffby\ufeffthe\ufeffthrust\ufefflevel\ufeffis\ufeffused\ufefffor\ufeffproviding\ufeffmotivation\ufefffor\ufeffthe\ufeffship.\ufeffThe\ufeff crew\ufeffadjusts\ufeffpropeller\ufeffspeed\ufeffvia\ufeffthrust\ufeffrod\ufeffso\ufeffas\ufeffto\ufeffcontrol\ufeffthe\ufeffforward\ufeffimpetus\ufeffof\ufeffthe\ufeffship.\ufeffIn\ufeffthis\ufeff paper,\ufeffthe\ufeffresponse\ufeffmodel\ufeffof\ufeffthe\ufeffpropulsion\ufeffunit\ufeffis\ufeffas\ufefffollows:", + "P\nR P P\nP P\nR P P\nX P a\nP a d\na d\nP a d\na = + <\n= \u2212 >\n\n\n \n= , ,\n,\n,0 \u03c4 \ufeff (1)\nwhere\ufeffP a \ufeffis\ufeffthe\ufeffactual\ufeffposition\ufeffof\ufeffthe\ufeffthrust\ufefflevel,\ufeffP d \ufeffis\ufeffthe\ufeffcommand\ufeffposition\ufeffof\ufeffthe\ufeffthrust\ufefflevel,\ufeff R P\n\ufeffis\ufeffthe\ufeffresponse\ufeffrate\ufeffof\ufeffthe\ufeffpropulsion\ufeffunit,\ufeffX\ufeffis\ufeffactual\ufeffpushing\ufeffforce,\ufeff\u03c4\ufeffis\ufeffthe\ufeffratio\ufeffcoefficient. Normalized\ufeffprocessing\ufeffof\ufeffequation\ufeff(1)\ufeffis\ufeffas\ufefffollows:\nX\nT X P\nX P\nT X P\nP d\nd\nP d\n'\n/ , '\n, '\n/ , '\n'\n'\n'\n=\n+ <\n=\n\u2212 >\n\n\n 2 0 2\n\ufeff (2)\nwhere\ufeff \u2032 = =X X X P P a / / max max ,\ufeffP P P d d ' max /= ,\ufeffT P R P P = 2 max / ,\ufeffP max \ufeffis\ufeffthe\ufeffmaximum\ufeff position\ufeffof\ufeffthe\ufeffthrust\ufefflevel,\ufeffX max \ufeffis\ufeffthe\ufeffmaximum\ufeffpushing\ufeffforce. Response Model of Steering Gear The\ufeffsteering\ufeffgear\ufeffcan\ufeffadjust\ufeffrudder\ufeffblade\ufeffposition\ufeffaccording\ufeffto\ufeffthe\ufeffcommand\ufeffrudder\ufeffangle\ufeffso\ufeffas\ufeffto\ufeff provide\ufeffthe\ufeffship\ufeffyawing\ufeffin\ufeffthe\ufeffyaw\ufefftorque.\ufeffThe\ufeffresponse\ufeffmodel\ufeffof\ufeffsteering\ufeffgear\ufeffin\ufeffthis\ufeffpaper\ufeffis\ufeffa\ufeff simplified\ufeffsteering\ufeffgear\ufeffas\ufeffshown\ufeffin\ufeffthe\ufefffollowing:\n\u03b4 \u03b4 \u03b4 \u03b4\n\u03b4 \u03b4 \u03b4 \u03b4 \u03b4\n\u03b4\n\u03b4\na\na d\na d\na d\nT\nT\n= + < +\n= + \u2212 > +\n\n\n 2 0 2 max max / , , / ,\n\u2206 \u2206 \u2206 \ufeff (3)\nwhere\ufeff \u03b4 a \ufeffis\ufeffthe\ufeffactual\ufeffrudder\ufeffangle;\ufeff \u03b4 max \ufeffis\ufeffthe\ufeffmaximum\ufeffrudder\ufeffangle,\ufeffusually\ufeff \u03b4 max = \u00b035 ;\ufeffT\u03b4 \ufeff is\ufeffthe\ufefftime\ufeffconstant\ufeffof\ufeffsteering\ufeffgear\ufeffresponse,\ufeffwhich\ufeffpresents\ufeffthe\ufeffminimum\ufeffduration\ufefffrom\ufeffa\ufeffhard\ufeffport\ufeff to\ufeffhard\ufeffright;\ufeff\u03b4\nd \ufeffis\ufeffthe\ufeffcommand\ufeffrudder\ufeffangle;\ufeff\u0394\ufeffis\ufeffthe\ufeffsteering\ufefferror\ufeffvector.\nResponse Model of Surging Ship\u2019s\ufeffsurging\ufeffis\ufeffalong\ufeffthe\ufeffx\ufeffaxis\ufeffin\ufeffthe\ufeffbody\ufeffcoordinate\ufeffsystem.\ufeffThe\ufeffresponse\ufeffmodel\ufeffof\ufeffsurging\ufeffshows\ufeff the\ufeffchange\ufeffrule\ufeffof\ufeffship\ufeffsurging\ufeffin\ufeffthe\ufeffform\ufeffof\ufeffdifferential\ufeffequations.\ufeffThis\ufeffmodel\ufefftakes\ufeffpushing\ufeffforce\ufeff and\ufeffhydrodynamic\ufeffresistance\ufeffin\ufeffto\ufeffaccount\ufeffrespectively\ufeffand\ufeffgets\ufeffthe\ufeffequation\ufeffof\ufeffsurging\ufeffacceleration\ufeff as\ufefffollows:\nM u X M r R u u u u = + \u2212\u03bd \ufeff (4)\nwhere\ufeffM u \ufeffis\ufeffthe\ufeffsum\ufeffof\ufeffship\ufeffweight\ufeffand\ufeffadditional\ufeffweight;\ufeffu\ufeffis\ufeffthe\ufeffspeed\ufeffof\ufeffadvance;\ufeffv\ufeffis\ufeffthe\ufeffswaying\ufeff speed;\ufeffr\ufeffis\ufeffship\ufeffyawing\ufeffspeed;\ufeffR\nu \ufeffis\ufeffhydrodynamic\ufeffresistance\ufeffcoefficient.\nWhen\ufeffthe\ufeffship\ufeffis\ufeffsailing\ufeffstably\ufeffwith\ufeffits\ufeffmaximum\ufeffspeed,\ufeffthe\ufefffollowing\ufeffequation\ufeffis\ufeffset\ufeffup:" + ] + }, + { + "image_filename": "designv11_92_0003117_978-3-030-01382-0_6-Figure6.14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003117_978-3-030-01382-0_6-Figure6.14-1.png", + "caption": "Fig. 6.14 Problem geometry and set of initial conditions", + "texts": [ + " In a closed-loop system, d becomes a function of some of the system\u2019s measured feedback sensors signals that are measured in real time. This is the feedback law. The closed-loop system then operates autonomously\u2014it is driven only by the system\u2019s initial conditions. This concludes Problem Formulation, and the related modeling theory. Each semester, slightly different initial conditions or parameters such as plane velocity were assigned. Solution (Benjamin Coleman, Fall 2016, Florida Atlantic University) The problem geometry is redrawn in Fig. 6.14. The block diagram is as shown in Fig. 6.15. Initial conditions R0, e0, and w0 are shown in blue. Display blocks (used for controlling and plotting the output of the simulation) are shown in purple\u2014these simulation blocks could be removed without changing the mathematic behavior of the system. The design parameters are shown in salmon. To estimate how long the plane will take to go from R = R0 to R = RF = 2 km, we can approximate the planes trajectory using straight lines. The quickest path will be a straight line directly towards the destination" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002529_012004-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002529_012004-Figure2-1.png", + "caption": "Figure 2. Illustration of mesh refinement", + "texts": [ + " Tools like VTMS and a very similar software called DFMA provide a method for producing non-idealized textile geometries, which is a step in the right direction, but significant work remains to ensure that the geometries closely resemble physical specimens. This work uses a standard FEA formulation, which requires a conforming mesh. The VTMS software includes meshing for the independent mesh method [12], but the resulting mesh is not usable herein because it is nonconforming. Instead, an in-house tool was used to create refined surface meshes of the tow geometries. The surface meshes were then used to create volume meshes of the tows and matrix using quadratic tetrahedrals, leveraging a general tetrahedral meshing library called TetGen [13]. Figure 2 illustrates the typical mesh refinement. Tows generally have 5 to 10 quadratic tetrahedrals through the thickness of the tows. A higher mesh refinement exists near binder tows. The final mesh used for this paper consists of 20 million nodes and 15 million quadratic tetrahedrals. It should be noted that the surface geometry that is taken from VTMS is faceted, resulting in unrealistic sharp corners between facets. Ideally, the faceted geometry should be smoothed to avoid unrealistic stress concentrations near sharp corners, but this is not done in this study" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001036_robio.2009.5420553-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001036_robio.2009.5420553-Figure4-1.png", + "caption": "Fig. 4. Schematic diagram of the level-adjustment device\u2019s actuators", + "texts": [ + " This is why two limit switches are installed at two ends of each linear motion unit to detect whether the linear motion unit moves to its limited position or not. Once it moves to one end, the limit switch will turn on and the linear motion unit should stop. Otherwise it is doomed to damage itself if it tries to move to the direction any further. Consequently, the industrial PC has to judge whether the linear motion unit can move further or not by detecting the switches\u2019 states. Two actuators of the device in a diagonal are schematically shown in Fig. 4. As the actuators\u2019 moving ranges are limited, the control strategy has to be modified when the linear motion unit reaches its limit position. For example, if x is positive, we should shorten 1R and loosen 3R to regulate the payload to level in the normal situation. However, 1R cannot be shortened if the limit switch #1 turns on, which means that the linear motion unit can only move to the opposite direction. In the situation, we can only loosen 3R to regulate x to zero, which will render 1R and 3R bear less tension" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002894_ihmsc.2018.10168-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002894_ihmsc.2018.10168-Figure3-1.png", + "caption": "Figure 3. simplified model of pitch axis", + "texts": [ + " the height axis is modeled in general, and the following dynamic differential equations are established: (2) Among them, J \u03b5 as the height of shaft inertia; \u03b5 high shaft deflection angle; K f motor constant; La body length to height of pivot points; p for the pitch axis deflection angle; V f andV b respectively before and after the voltage of the motor for the components of gravity. 270 978-1-5386-5836-9/18/$31.00 \u00a92018 IEEE DOI 10.1109/IHMSC.2018.10168 ( )bffa VVKpLJ += \u2022\u2022 cos\u03b5\u03b5 3 B. Pitching Axis Modeling According to the simplified model shown in Figure 3, the pitch axis movement by the difference between the two produced by the propeller thrust control, in the case that the current to the motor thrust is greater than the motor thrust, pitching axis Angle is positive, the helicopter fuselage would be positive pitch model. The dynamic model of the helicopter in the pitching axis is as follows: ( )bffhp VVKpLJ p \u2212= \u2022\u2022 cos 4 Among them, J p the moment of inertia for the pitch axis; The deflection Angle of pitch axis p ; K f Is the motor constant; The length of the pitch axis to the propeller Lh " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003010_052022-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003010_052022-Figure1-1.png", + "caption": "Figure 1. ABB IRB 140 link dimensions [2]", + "texts": [ + " Before being installed in the faculty lab, this robot was mainly used as a test and training robot. Over time (since 2007), the robot has been transported tens of times to different locations where tests or trainings have been done. He suffered strikes and was serviced multiple times. From the information that can be read on his teachpendant, 986 hours have passed since the last service from which 658 hours of effective operation. The last work done on it was changing of the discharged batteries (1 week before the measurements). More robot specification can be found in Table 1 and Figure 1 below). 2 1234567890\u2018\u2019\u201c\u201d IOP Conf. Series: Materials Science and Engineering 444 (2 18) 052022 doi:10.1088/1757-899X/444/5/052022 In order to evaluate robot accuracy and repeatability a laser tracker measurement system was used. The measurement system is a RADIAN API laser tracker [7]. This laser-tracker system is a dynamic (continuous acquisition) measurement system that uses spherical mirror retro-reflectors (SMRs) to track the laser beam. It is based on a laser interferometer (IFM) system with the ability to measure unparalleled distances and has an \u201con-board reference\u201d measurement technology" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000120_indin.2008.4618144-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000120_indin.2008.4618144-Figure2-1.png", + "caption": "Fig. 2. The assignment of the coordinate system", + "texts": [ + " Secondly, the transformation between the robot tool and the external axle worktable is calibrated by the robot and the contact position sensor joint measurement system. Finally, Particle Swarm Optimization (PSO) Algorithm is adopted to optimize the key parameters to improve the accuracy of the total system. The robot scanning system with an external axle is shown in Fig.1. The system consists of an external axle and a six-degree freedom robot with an inductance contact position sensor fixed on its wrist flange. The robot coordinates is shown in left part of Fig.2. The frame {B}, located on the center of the robot\u2019s base, is the reference coordinates of the system. The center of the flange at the bottom of the robot is defined as frame {W}. The zero reading point of the sensor is defined as the 1 The work was supported by national natural science foundation of China (50575029) and (60674061) 2 Corresponding author: e-mail: zima@newmail.dlmu.edu.cn \u24d2 origin of the frame {T}. The right part in Fig.2 shows an external axle with two degrees of freedom (rotating and pitching). Frame {0}, defined as the reference frame, is fixed on the pitching axis and aligns with frame {1} when the first joint variable ( 1) is zero. The Z-axis of frame {0} and {1}, called Z0 and Z1, is coincident with the pitching axis. The origins of frame {0} and {1} are the same point, which is located where the perpendicular a1 intersects the joint rotating axis, so a0 is the distance from Z0 to Z1 , 0 is the angle between Z0 and Z1 and di is the distance from X0 to X1 are zero", + " The public perpendicular intersecting point in rotating axis is the origin of frame {2}. Z2 is coincident with the rotating axis. X2 is coincident with a1. 2 is the angle of rotating axis. The Frame {2} is the worktable frame. The link parameters of the external axle are listed in table I. i i-1 ai-1 di i 1 0 0 0 1 2 1 a1 0 2 The calibration from wrist to sensor can determine the position of the sensor to the robot wrist flange. The Relationship between the different coordinates system {T}, {W} and {B} shown in Fig. 2 can be expressed by a congruence transformation of the form 1 1 0 1 0 1 B W B TW BW T WTP PR T R T = ( 1 ) Where PB represents an point in frame {B},and B w R describes the orthogonal rotational matrix of frame {W} relative to {B}, and TBW expresses the translation vector from {B} to {W}, and W T R is the orthogonal rotational matrix of frame {T} relative to {W}, and TWT is the translation vector from {W} to {T}. Because the sensor is fixed on the center of the wrist flange, W T R and TWT can be calibrated directly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000787_mfi.2008.4648048-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000787_mfi.2008.4648048-Figure2-1.png", + "caption": "Fig. 2. Structure of 5-link biped model", + "texts": [ + " Therefore, the most of the biped robot have been modeled as the 5-link biped model. In order to realize a complex gait of human on the simplified biped robot, the dynamics of the human have to be associated with its one. Modeling the dynamics of the human gait have been worked by lots of physiologists. They make it show that the dynamics of human gait takes place on the sagittal plane, or the plane bisecting the human body as shown in Fig.1. [3] I The 5-link biped model is constructed as shown in Fig. 2. The biped model above consists of five rigid links, one link for trunk, two links for thigh, and two links for shank. This 5-link biped model parameters are represented as follows: id : the distance from joint i to COM of link i iI : the inertia moment of link i i\u03b8 : the angle of link i ( , )e ex y : the coordinate of supporting point ( , )b bx y : the coordinate of the tip of the swing limb i\u03c4 : the torque of link i For measuring these parameters, a subject puts on the helpful walking device and attaches four markers to his head, pelvis, knee, and ankle as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000347_peds.2009.5385847-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000347_peds.2009.5385847-Figure11-1.png", + "caption": "Fig. 11. Stator flux locus during acceleration speed from 0.5 to 2.5 p.u. for (a) DTC without overmodulation (b) DTC with the proposed overmodulation", + "texts": [ + " Thus a shorter acceleration time is obtained during the speed transient response with the proposed overmodulation method. In general, the overmodulation mode in DTC produces higher current harmonics content as a result of the hexagonal shape of the stator flux. Apparently, the stator current distortions only occur when the proposed overmodulation method is applied in DTC to achieve higher torque capability and hence rotor to accelerate faster. The PWM mode of stator voltage and current will be operated again as the speed reaches its reference. Fig. 11, shows the stator flux locus corresponds to the results obtained in Fig. 10. In the proposed overmodulation method, the stator flux locus is suddenly changed to the hexagonal shape. In such manner, the stator flux angular velocity increases as the hexagonal stator flux locus shrinks (flux weakening) as shown in Fig. 11. This paper presents a simple overmodulation method employed in DTC-CSF based scheme. It is shown that, the stator voltage can perform under overmodulation mode by changing the stator flux locus from circular to the hexagonal shape. An extension of constant torque region and higher capability of torque in field weakening region can be obtained with the proposed overmodulation method. The main benefit in the proposed strategy is its simplicity and it can be implemented in many electric drive applications to obtain maximum torque capability for a wider speed range operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001813_978-981-10-8306-8_13-Figure13.8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001813_978-981-10-8306-8_13-Figure13.8-1.png", + "caption": "Fig. 13.8 The p diagrams of the Hosford set and reference states: a yield surface and b plastic strain increment surface", + "texts": [], + "surrounding_texts": [ + "Derive for Eq. (13.44) that b \u00bc 2M 2\u00fe 2M or b \u00bc 2M 1, respectively, when the reference state is simple tension (and balanced biaxial) or pure shear, while, dY \u00bc dB \u00bc \u00f01\u00fe 2M 1\u00de 1 M dK, considering Eqs. (13.27) and (13.29). Note that YdY = BdB = KdK, regardless of the M value, considering the relationship between Y, B and K of the Hosford yield function shown in HW #12.13. In fact, the Hosford set are conjugates of each other for those reference states. Prove this, considering the dual normality rules." + ] + }, + { + "image_filename": "designv11_92_0001065_icmtma.2009.276-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001065_icmtma.2009.276-Figure2-1.png", + "caption": "Figure 2. Whole finite-element meshes for the compressor", + "texts": [], + "surrounding_texts": [ + "(1) (2)\n(1) (2)\nif 0 on if 0 f n c\nf n\np p du du S\np p du du\n\u03c4 \u03c4 \u03c4\n\u03c4 \u03c4 \u03c4\n\u03bc\n\u03bc\n< \u2212 \u2212 = = \u2212 \u2212 \u2265 (8)\nwhere the physical meaning of all the parameters and variables is identical with the conventional one, which can also be found in Refs. [5].\nThe friction constraints of 3D contact problem are different from the ones of two-dimensional problem and make the contact problem more difficult to handle. It can be described by introducing a closed set formulation of friction law\n( ) { }: 0n f nC p f p p\u03c4 \u03c4 \u03bc\u221e = = + \u2264P The graph of Coulomb\u2019s friction law is a quadratic cone. In order to get the LCP formulation, a piecewise linearization of Coulomb\u2019s friction law is introduced\n( ){ } [ ]1 2 : , 0, 1, 2, ..., cos , sin , , , fN i n f T i i i f n C f p p i N f p p p \u03c4 \u03c4 \u03c4 \u03c4\u03b1 \u03b1 \u03bc = \u2264 = = P (9)\nIntroducing penalty factors , n\u03c4E E in tangential and normal directions of the contact surface, the contact strains is then defined as\ne p c c cd d d= +\u03b5 \u03b5 \u03b5 (10)\n( )p c c c cd d d= \u2212\u03b5 \u03b5p D (11)\nwhere\n[ ] [ ] 1 2\n1 2\n(1) (2) 1 1 1 1\n(1) (2) 2 2 2 2\n(1) (2) * *\n, ,\n, ,\n0 0 0 0 0 0\nT c n\nT c n\nc\nn\nun n n n\nd d d d\nd dp dp dp\nd du du u\nd du du u\nd du d u\n\u03c4 \u03c4\n\u03c4 \u03c4\n\u03c4\n\u03c4\n\u03c4 \u03c4 \u03c4 \u03c4\n\u03c4 \u03c4 \u03c4 \u03c4\n\u03b5 \u03b5 \u03b5\n\u03b5\n\u03b5\n\u03b5 \u03b4 \u03b4\n=\n=\n=\n= \u2212 = \u0394\n= \u2212 = \u0394\n= \u2212 + = \u0394 +\np\nE D E\nE\n\u03b5\n, (12)\nand *\u03b4 is the initial gap. Defining sliding \u201cyield\u201d function\nif and \u201cflow\u201d potential function ig along the contact surfaces as\n1 2\n1 2\n1\n1\ncos sin 0,\ncos sin ,\n0,\n, 1, 2,...,\nf\nf\ni i i f n\ni i i\nN n\nN n\nf p p p\ng p p c\nf p\ng p i N\n\u03c4 \u03c4\n\u03c4 \u03c4\n\u03b1 \u03b1 \u03bc\n\u03b1 \u03b1\n+\n+\n= + + \u2264\n= + \u2212\n= \u2264\n= =\n, (13)\nthen the sliding strain can be expressed as\n1\n1 , 0\nf\nc\nT N\np i c i\ni c\nd \u2202 \u2202 \u2202 \u2202\n+\n= = = \u2265 g g p p \u03bb \u03bb \u03bb , (14)\n\u03bb is called slip factor vector or parametric variable vector and the boundary/complementary conditions on cS\ncan be written as the following form\n0 0\n0, 0, 0,\n1, 2, , 1f\ni i c i i\ni i i i\nf d\ni N\n+ \u2212 + =\n\u2265 \u2265 \u22c5 =\n= +\nW M v\nv v\n\u03b5 \u03bb\n\u03bb \u03bb (15)\nwhere 0 i f is the initial value of if , and\n,\nT T T\ni i i c c\nc c c\nf f\u2202 \u2202 \u2202 \u2202 \u2202 \u2202 = = g W D M D p p p\n,\nThe parametric variational principle [5, 6] for 3D frictional contact problems can be established based on the previous discussion: among all admissible displacement incremental fields, in which satisfy the strain-displacement relations and geometrical boundary conditions, the actual displacement field taken by a body is the one that makes the parametric total potential energy [ ]\u03a0 \u03bb minimum under the control of status equations (11), i.e.,\n[ ] , , 1\n( ) 2\n1\n2c\nP\ni j ijkl k l i i\nci cij ij k ck S\ni i S\nd d d d d\nd d d ds\nd d ds\n\u03b1 \u03b1\u03bb\n\u03a9 \u03a0 \u22c5 = \u2212 \u03a9\n+ \u2212\n\u2212\n\u03bb\n\u03b5 \u03b5 \u03b5\nu D u b u\nD R\np u\n, (16)\nwhere\n[ ] [ ] [ ] [ ] 1 2 3 1 2 3\n1 2 3 1 2 3\n, , , ,\n, , , ,\nk cik ci\nT T c c c c\nc c c c\ncij ij ij n ij\nd d d d d d d\nT T d dp dp dp dp dp dp\n\u03b1 \u03b1\n\u03c4 \u03c4 \u03c4\n\u03c4 \u03c4 \u03c4\n\u03c4\n\u2202 \u2202\n\u03b5 \u03b5 \u03b5 \u03b5 \u03b5 \u03b5\n\u03b4 \u03b4 \u03b4\n=\n= =\n= =\n= = =\n\u03b5\ng R D\np\np\nD E E E\nIt can be proved that PVP may be changed into a quadratic programming problem through derivation. And a normal quadratic programming problem is a convex programming problem, which has mature solution method (e.g. Lemke algorithm) and can ensure high efficiency in computation.\nIII. DETAILS OF THE COMPUTATIONAL MODEL The compressor impeller is made up of identical segments in the circumferential direction and hence constitutes a rotationally periodic structure. The impeller with a shaft hole 34mm in diameter is mounted onto the sleeve and the compressor shaft via interference fit.", + "The material of the sleeve and shaft is steel and the material of the impeller is aluminum. The computation needs to be carried out by using 3D model, instead of by axial symmetric one. The computational model of substructure and the impeller is shown in Figs. 2. The model of substructure is shown in Fig 3.\nThe main loads applied to the compressor are interference-fit load and centrifugal forces caused by highspeed rotation when operating. 1\u03b4 is defined as the amount of interference between impeller and the shaft sleeve and\n2\u03b4 as the amount of interference between the shaft sleeve\nand shaft. For different values of 1\u03b4 =0.04 mm and\n2\u03b4 =0.05mm, the load cases are computed respectively for various rotational speed n (0, 24000, 25000, 27000, 29000rpm), various coefficients of friction \u03bc (0.1, 0.15, 0.2) and various wall thicknesses t (7, 6, 5, 4 mm).\nIV. RESULTS AND DISCUSSIONS\nA. The influence of the rotational speed In Figs. 4~7, solid curves represent contact stress distributions on the external surface of shaft sleeve (surface between impeller and shaft sleeve) and dashed curves represent the contact stress distributions on the internal surface of the shaft sleeve (the surface between the shaft sleeve and the shaft).\nContact stresses distribution of both the external and internal surface of shaft sleeve along the axial direction are illustrated in Fig. 4 ( 1\u03b4 =0.04mm, 2\u03b4 =0.05mm) with four rotational speeds. In Fig. 4, it can be seen that for the same amount of interference, contact stress of mating surface decreases with the increase of rotational speed due to the increase of centrifugal forces. Variation of the contact stress of mating surface edge is small and the change trend of the contact stress of internal and external surface is uniform. Variation of the contact stress of the external surface edge of shaft sleeve is small owing to the small blade mass at the external surface edge of shaft sleeve.\nFor x=8.375mm, computed results show that relative slips along the tangent direction of shaft occur on the external surface of shaft sleeve (see Fig. 5). With the increase of the amount of interference between impeller and shaft sleeve, the slippage of sliding point increases linearly.\nB. The influence of friction coefficient The coefficient of friction between impeller and shaft sleeve is case 1. \u03bc =0.1; case 2. \u03bc =0.17; case 3. \u03bc =0.2. From Fig. 6 it is clear that with the increase of friction coefficient between impeller and shaft sleeve, the contact", + "stress of mating surface keeps unchanged. Lubricating oil smeared on mating surface has effect on decreasing contact stress by press-fitting, but mating surface can prevent from block and abrasion by using lubricating oil. So the pure vegetable oil is used as lubricant during press-fitting.\nC. Influence of the Wall Thickness of Shaft Sleeve For the central contact point x = 21.25mm, the relation of rotational speed and the contact stress is given for different wall thickness of shaft sleeve. In Fig. 7, solid curves represent the contact stress distribution of the external surface of shaft sleeve; dashed curves represent the contact stress distribution of the internal surface of shaft sleeve. With the increase of the rotational speed, the contact stresses of both the external and internal surface of shaft sleeve descend. At the same rotational speed, the contact stress of the external surface of shaft sleeve decreases with increase of the wall thickness of shaft sleeve, but the contact stress of internal surface increases instead. It is the reason that the wall thickness of shaft sleeve varies by changing radius of the external surface of shaft sleeve in computational model, the wall thickness of shaft sleeve increases means the internal diameter of impeller increases, the rotational mass of impeller decreases when the amount of interference between the shaft sleeve and shaft remains unchanged. Consequently, the contact stress of the external surface of the shaft sleeve decreases. In the same way, the reason that the contact stress of the internal surface of the shaft sleeve increases can be explained.\nV. CONCLUSIONS The turbocharger compressor three-dimensional frictional contact problem using the FE PQP method and substructure technique is analysed. The computed results obtained provide an effective approach to improve design and manufacturing technology of compressor impellers. At the rated speed of the compressor, the relative displacement of the corresponding contact region occurs at the region where the radial size of the impeller is larger, and 30% contact points between the impeller and shaft sleeve may be in disengaged positions. Relative slip along the tangential direction occurs on the external surface of shaft sleeve have a bad influence on fatigue resistances of mating surfaces. The influence of friction coefficient is small to axial variation of the contact stresses. With the increase of wall thickness of shaft sleeve, contact stress of external surface of shaft sleeve decrease and the change trend of internal surface increase within the certain range of the rotational speed. To maintain steady and enough contact stress, the wall thickness of shaft sleeve should adopt smaller value to avoid the rapid change of the contact stress along the axial direction.\nThis work is jointly supported by Leading Academic Discipline Project of Shanghai Municipal Education Commission, Project Number: J51401.\n[1] D. V. Bhope and P. M. Padole, \u201cExperimental and theoretical\nanalysis of stresses, noise and flow in centrifugal fan impeller\u201d, Mechanism & Machine Theory, vol. 39, Dec. 2004, pp. 1257-1271, doi:10.1016/j.mechmachtheory.2004.05.015.\n[2] H. W. Zhang, W. X. Zhong and Y. X. Gu, \u201cA combined parametric quadratic programming and iteration method for 3D elastic-plastic frictional contact problem analysis\u201d, Computer Methods in Applied Mechanics and Engineering, vol. 155, Mar. 1998, pp. 307-324, doi:10.1016/S0045-7825(97)00170-9.\n[3] H. W. Zhang, S. Y. He, X. S. Li and P. Wriggers, \u201cA new algorithm for numerical solution of 3D elastoplastic contact problems with orthotropic friction law\u201d, Computational Mechanics, vol. 34, Jun. 2004, pp. 1-14, doi: 10.1007/s00466-004-0548-2.\n[4] H. W. Zhang, A. H. Liao and C. H. Wu, \u201cSome Advances in Mathematical Programming Method for Numerical Simulation of Contact Problems\u201d, IUTAM Symposium on Computational Methods in Contact Mechanics, Springer Press, Nov. 2007, pp. 33-55, doi:10.1007/978-1-4020-6405-0_3." + ] + }, + { + "image_filename": "designv11_92_0002552_012026-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002552_012026-Figure1-1.png", + "caption": "Figure 1. The carrier position on inclined plane scheme", + "texts": [ + " Then all the forces acting on the system: ~PC \u2013 the carrier weight, ~PL \u2013 the load weight, ~N \u2013 the sloped inclined plane normal reaction and ~Ffr \u2013 the sliding friction force in rest are situated in the plane Q. Since ~PC \u2191\u2191 ~PL and, hence, ( ~PC , ~PL ) \u223c ~P , where ~P = ~PC + ~PL, then the carrier will be in rest on the inclined plane if the system of forces acting on it is balanced, i.e. ( ~P , ~N, ~Ffr ) \u223c 0. As the considered system is a system of convergent forces on plane, then writing its balance conditions we obtain (see fig. 1) 3\u2211 s=1 Fsx \u2261 \u2212 (PC + PL) \u00b7 sin\u03b1+ Ffr = 0 , (1) 3\u2211 s=1 Fsz \u2261 \u2212 (PC + PL) \u00b7 cos\u03b1+N = 0 . (2) Let\u2019s add to (1), (2) the Coulomb-Amonton law Ffr \u2264 Ffrmax , (3) where Ffrmax = f \u00b7N and f is the coefficient of sliding friction in rest for the pair of materials \u201ccarrier-inclined plane\u201d. The conditions (1), (2) are the balance equations. Solving them we can find Ffr and N . The substitution of Ffr and N into (3) allows obtain the condition for the plane inclination angle tg\u03b1 \u2264 f . Further on the condition is presented as follows |\u03b1| \u2264 arctg f . (4) 3 1234567890 \u2018\u2019\u201c\u201d IOP Conf. Series: Journal of Physics: Conf. Series 973 (2018) 012026 doi :10.1088/1742-6596/973/1/012026 Note. On fig. 1 HH is the line of intersection of the horizontal plane and the vertical plane Q; the plane inclination angle \u03b1 > 0. According to [6] let\u2019s introduce: Oxyz is the motionless coordinates system being an inertial counting out system, where Oxy is the plane coinciding with the inclined plane; the axis Ox is directed along the line of the greatest descent (to ascent for \u03b1 > 0 or to descent for \u03b1 < 0), and axis Oz is directed normally to inclined plane; O1x1y1z1 is the mobile coordinates system fastened together with the carrier, besides, O1x1z1 coincides with the vertical plane, and O1x1 \u2191\u2191 Ox ; O2x2y2z2 is the mobile coordinates system fastened together with the channel, besides, O2x2z2 coincides with the vertical plane, and \u03d5 is the channel installation angle \u2013 the angle between axes O1x1 and O2x2 counting up from inclined plane for \u03d5 > 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003018_ilt-07-2018-0272-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003018_ilt-07-2018-0272-Figure7-1.png", + "caption": "Figure 7 Run-in surfaces of a gear contact at point B on the path of contact", + "texts": [ + " Several surface sections of various teeth are considered owing to the limited size of the measured surface section to obtain a Figure 6 Scheme TEHL simulation and pit lifespan calculation from rough concentrated contacts Tooth flank load carrying capacity Martin Zimmer and Dirk Bartel Industrial Lubrication and Tribology D ow nl oa de d by Y or k U ni ve rs ity A t 1 5: 36 1 7 D ec em be r 20 18 ( PT ) better statistic base for the TEHL simulation and the stress calculation. Two run-in surfaces that are used for the estimation of three-dimensional (3D) micro stress fields and theTEHL simulation are shown in Figure 7. For these surfaces, the integral solid body contact pressure curves and the pressure and shear flow factors are calculated. By doing so, the effects of the rough surfaces on the load capacity of the lubricant film and the formation of the lubricating gap are determined, thus enabling the calculation of the friction condition. Taking into consideration the temperature development in the lubricating gap, the pressure and shear stress distribution in the concentrated contact for the respective operating conditions can be calculated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002529_012004-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002529_012004-Figure9-1.png", + "caption": "Figure 9. Contours of each stress component in the local coordinate system normalized by the respective strength for the case of uniaxial tension along the global y-axis", + "texts": [ + " In this region, in the weft drops to about 0, such as point C in Figure 8a. The drop in transverse tension is due to a compressive (global) in the region where a binder crosses the wefts at the top or bottom of the textile, such as region D highlighted in Figure 8b. The compressive stress forms as load is transferred to or from the binder via shear, since the binder is orders of magnitude stiffer in the global x-direction than the surrounding material at the top and bottom. 3.2.2 Tension Along Y-Axis Figure 9 shows three most severe normalized stresses within the clipped analysis region for an applied volume average strain along the global y-axis, , of 1%. Again, the normalized stresses shown are in the local coordinate system. Overall, Figure 9 shows that the warps and binders experience severe transverse tension ( ), while the wefts experience severe longitudinal shear ( and ). The most severe stress in the textile is in the binders and warps. For both types of tows, the local y-axis is closely aligned with the global y-axis, which is the direction of the load for this configuration. To illustrate where the peak stresses occur in the textile, Figure 10 shows for one binder and row of warp tows, along with the local coordinate system for a point in the binder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000742_6.2009-2404-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000742_6.2009-2404-Figure1-1.png", + "caption": "Figure 1: Generic HALE UAV", + "texts": [ + " In the remainder, we will use the subscript r to denote the quantities associated with the right-half wing, l with the left-half wing, f with the fuselage, h with the horizontal stabilizer and, finally, v with the vertical stabilizer. The motion of the aircraft can be conveniently described by attaching a set of body axes xyz to the undeformed aircraft at a convenient point, as well as similar body axes to the left- and the right-half wing, namely, xryrzr and xlylzl with origins or and ol, respectively (Fig. 1). The motion can be expressed by three translations and three rotations of the body axes xyz and elastic deformations of the left- and the right-half wing relative to the respective body axes. The generic equations of motion of the whole aircraft from Ref. 7 are d dt ( \u2202L \u2202v ) + \u03c9\u0303 \u2202L \u2202v = F d dt ( \u2202L \u2202\u03c9 ) + v\u0303 \u2202L \u2202v + \u03c9\u0303 \u2202L \u2202\u03c9 = M (1) \u2202 \u2202t ( \u2202L\u0302 \u2202vi ) \u2212 \u2202L\u0302 \u2202ui + \u2202F\u0302i \u2202u\u0307i + Liui = U\u0302i \u2202 \u2202t ( \u2202L\u0302 \u2202\u03b1i ) + \u2202J\u0302i \u2202\u03c8\u0307i + Hi\u03c8i = \u03a8\u0302i, i = r, l where L = T \u2212V is the Lagrangian for the whole aircraft, in which T is the kinetic energy and V is the potential energy, v = [U V W ]T and \u03c9 = [P Q R]T are the vectors of translational and angular velocities of xyz, C matrix of direction cosines between xyz and inertial axes XY Z, R = [X Y Z]T position vector of the origin o of xyz relative to XY Z, E matrix relating Eulerian velocities to angular quasi-velocities, \u03b8 = [\u03c6 \u03b8 \u03c8]T symbolic vector of Eulerian angles between xyz and XY Z, ui = [0 vi wi]T and vi = u\u0307i elastic bending displacement and velocity vectors, \u03c8i = [\u03d5i 0 0]T and \u03b1i elastic torsional displacement and velocity vectors, L\u0302 Lagrangian density exclusive of strain energy, F\u0302i and J\u0302i Rayleigh\u2019s dissipation function densities, Li and Hi matrices of stiffness differential operators, F and M resultant of gravity, aerodynamic, propulsion and control force and moment vectors acting on the whole aircraft in terms of body axes components, U\u0302i and \u03a8i resultant of gravity, aerodynamic, propulsion and control force and moment density vectors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure40-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure40-1.png", + "caption": "Fig. 40 Distributed sensors on the X-spoke-type strain structure", + "texts": [ + " We found CrN-STFs, which are CrN (nitrogen-added chromium) alloy thin films and have high strain sensitivity, could be applied and adapted into joint torque sensors. The CrN-STFs have been developed by Research Institute for Electromagnetic Materials [25]. The stiff design on flexure element of the joint torque sensors is provided by applying high sensitive and resistive thin-film strain sensors. Introducing the CrN-STFs in a new strain sensing method enables stable, reliable, and high-resolution torque sensing with high signal-to-noise ratio. Figure 40 shows the layout of the distributed strain sensors on the joint torque sensor. The joint torque sensor has an X-spoke-type strain structure comprised of inner race (fastener part), outer race (fastener part), and four strain beams. On each strain beam, a pair of resistive thin-film strain sensors (R1 R8) is located and two thermostatic thin-film sensors (thermosensors, R9 and R10) are located in two positions of non-deformable area. The strain sensors and thermosensors are formed on the insulating film (SiO2) applied to surface-polished stainless steel (SUS316L or SUS630) strain structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002873_smasis2018-7915-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002873_smasis2018-7915-Figure3-1.png", + "caption": "FIGURE 3: FUNCTIONAL MODEL FOR THE ACTIVE INFLUENCE ON THE CURVATURE OF THE BEARING SHELLS", + "texts": [ + " On the contrary, a neutral, stable whirl motion of the journal (mostly known as limit cycle) with a large amplitude is detected in the case of the circular bore for a long region of the investigated speed range, before the instability occurs. This observation can be interpreted as a result of significantly increasing the lubricant film stiffness and minimizing it\u2019s nonlinear behavior in the case of strongly disturbing the circularity of bore clearance, as far as the total bearing stiffness can be assumed as a serial set-up of structural and air film stiffnesses. Fig.3 shows a functional model for the AAFB (without mounting the foil structure) in a circular, non-active configuration, based on the concept explained in the last section. The functional model is made of stainless steel with an inside diameter of 70 mm, a width of 40 mm and thickness of 0.4 mm. Three Tough Resin (Formlabs) printed flexure joints are attached to the end of each support shell to allow the sliding movement of each connection point of the support shells in the radial direction. For actuation based on the inverse piezoelectric effect, three standard DuraAct P-876" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000615_paciia.2008.140-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000615_paciia.2008.140-Figure1-1.png", + "caption": "Fig. 1 3D model of the underwater manipulator", + "texts": [ + " In order to improve the tracking ability, a new compound control of method based on the cerebellar model articulation controller (CMAC) is proposed. The CMAC and the PID controller are connected in parallel. A nonlinear tracking differentiator (NTD) is presented to get high quality differential signals for the PID controller. Simulation experiments results show that the compound method is efficiency and has strong anti-interference. The underwater manipulator is a 7-function manipulator which is used primarily in underwater robot application. The 3D model of the manipulator is shown in Fig.1, and the structure is shown in Fig.2. The underwater manipulator has six DOF and a paw, including shoulder yaw, shoulder pitch, elbow pitch, forearm roll, wrist pitch and wrist roll. The operating depth is 3000 meters, and hydraulic drive is adopted because it is easy to seal in the underwater environment. The forearm roll joint is driven by swinging cylinder, while the wrist roll joint is driven by hydraulic motor, and the other joints are driven by hydraulic cylinder. 978-0-7695-3490-9/08 $25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000656_1.2992222-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000656_1.2992222-Figure1-1.png", + "caption": "Figure 1. A fabricated LbL assembled SWNT multilayer thin-film resistor: (a) Schematic diagram of multilayer hierarchy, (b) a photograph of the individual device, and (c) current-voltage (I-V) characteristics testing apparatus at pH buffers", + "texts": [ + " f-SWNT samples for FTIR were prepared by drop casting of SWNT solution onto Si/SiO2 surface cleaned as mentioned previously. The reflective Fourier transform infrared spectroscopy (Nicolet Magna IR 750) was used with the background of Si/SiO2 surface to characterize induced surface functional groups. The spectra were obtained as the absorbance in the midinfrared (MIR) region, 4000-400 cm-1 by averaging 1000 scans with a spectral resolution of 2 cm-1. A fabricated LbL assembled SWNT thin-film resistor is illustrated in Figure 1. In LbL multilayer hierarchy, as shown in Figure 1(a), there are (PDDA/PSS)2 as a precursor layer for the charge enhancement and (PDDA/SWNT)5 as an electrochemical transducer material. The spun PMMA in the outmost layer plays a role of the dielectric. The conductance behavior and pH sensing mechanism will be discussed in the absence and presence of the dielectric PMMA. The fabricated SWNT thin-film resistor depicted in Figure 1(b), was submerged into pH buffer solutions with a Ag/AgCl reference electrode which was connected to source electrode to provide a stable potential, as shown in Figure 1(c), and the current flowing through SWNT multilayer was acquired at the bias voltage of 0 to 1 V. The SEM images of the fabricated device are shown in Figure 2. The sourcedrain electrodes along with SWNT multilayer thin-film in the channel area are shown in Figure 2(a). The dimension of resistor channel used in this study is 10 \u00b5m long and 1 mm wide. Figure 2(b) illustrates the surface of (PDDA/PSS)2(PDDA/SWNT)5. The individual SWNT, its bundle, and random network are clearly observed. 5 ) unless CC License in place (see abstract)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002358_gt2018-75432-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002358_gt2018-75432-Figure1-1.png", + "caption": "Figure 1: Sectioning Plan used for the Study", + "texts": [ + "010 Table 2: Heat Treatments used for the Study Heat Treatment-1 Heat Treatment-2 Heat Treatment-3 Pre-weld HT 2050\u00b0F for 4 hours Pre-weld HT 2050\u00b0F for 4 hours with slow or fast cool HIP at 2190\u00b0F \u00b125\u00b0F for 4 hours with minimum 15K PSI in Argon Post-weld HT 2050\u00b0F for 2 hours Post-weld HT 2150\u00b0F for 2 hours Full solution at 2175\u00b0F \u00b125\u00b0F for 2 hours in vacuum and argon quenched Diffusion- MCrAlY coating 2050\u00b0F for 2 hours Diffusion- HVOF coating 2050\u00b0F for 2 hours partial solution and age at 2050\u00b0F \u00b125\u00b0F for 2 hours (controlled cooling) Diffuse TBC coating 2050\u00b0F for 2 hours Diffusion- TBC coating 2050\u00b0F for 2 hours Diffusion- HVOF coating 2050\u00b0F for 2 hours Age 1550\u00b0F for 24 hours all in vacuum Age 1550\u00b0F for 24 hours all in vacuum Diffuse TBC coating HT 2050\u00b0F for 2 hours 1550\u00b0F for 24 hours and argon quenched both in vacuum. The sectioning plan for microstructural investigations included five sections taken across the length of the blade as shown in Figure 1. The details of the sections are provided in Table 3. The root section is used for comparison of microstructure from the new manufacture, as this area is not exposed to higher temperatures. Microstructural investigations included low magnification optical microscopy to determine porosity levels, carbide structures and high magnification scanning electron microscopy to study and compare the gamma-prime morphology. Gamma prime size was measured using SEM image analysis provided with the EDAX software" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002196_acc.2018.8431927-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002196_acc.2018.8431927-Figure1-1.png", + "caption": "Fig. 1. Moving inverted pendulum stabilized by vibrational control method.", + "texts": [ + " It shows that the averaged trajectories converge arbitrarily close to the ultimate bound \u03b3\u03021(\u2016w\u2016\u221e). In the special case of w = 0, the averaged trajectories will converge arbitrarily close to the equilibrium point which means the system is v-stabilizable. If the equilibriums xe,v and ze satisfy condition (7), the original trajectories of the solutions converge to the ultimate bound. Vibrational control has been shown to be a useful method to stabilize the inverted pendulum without a feedback. The mechanism is shown in Figure 1. Assume a massless slider is connected to the suspension pin of an inverted pendulum by a revolution joint. The slider acts as an actuator to guide the vertical movement of suspension pin. It is shown in [5] that if the displacement of the slider is a sine wave with fast frequency, the inverted pendulum could be stabilized. Next, instead of stabilizing the pendulum to a still upper point, we assume the targeted upper point of pendulum is moving vertically at a constant speed. Thus besides the sinusoidal dither, an extra displacement signal \u03c3(t) = \u03c3t which casts the moving target will be added to the motion of the slider, where \u03c3 is the speed of the target" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003780_062005-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003780_062005-Figure5-1.png", + "caption": "Figure 5. The distribution diagram of static magnetic field.", + "texts": [ + " The air-gap flux density is distorted throughout the circumference. SCSET 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1176 (2019) 062005 IOP Publishing doi:10.1088/1742-6596/1176/6/062005 In order to analyse the air-gap magnetic field harmonic distribution after rotor eccentricity, FFT analysis of air-gap magnetic field harmonic distribution is given with an eccentricity of 25% on the condition of static and dynamic motor model, the harmonic content and harmonic amplitude of the airgap flux density are obtained as shown in Figure 5. SCSET 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1176 (2019) 062005 IOP Publishing doi:10.1088/1742-6596/1176/6/062005 As can be seen from Figure 5, the amplitudes of the original fundamental wave, 3,5,7,9,11th harmonics remain basically unchanged after the rotor eccentricity of the motor, but the amplitude of the harmonics with the \u00b1 f/p order is obviously increased, and the harmonic component in the air-gap magnetic field becomes larger. Radial electromagnetic force produced in air-gap magnetic field of motor radiates noise from the surface of stator and rotor to the external radiation noise caused by radial electromagnetic force acting on the surface of stator and rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001654_978-3-319-75118-4_7-Figure7.16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001654_978-3-319-75118-4_7-Figure7.16-1.png", + "caption": "Fig. 7.16 Schematic diagram of electron beam welding process", + "texts": [ + " \u2022 The high-intensity electron jet or beam passes through electromagnetic lens and defecting coil in order to focus the electron beam at the needed area. The high- velocity electron beams are directed to the weld cavity where it is used for the welding process. The high-kinetic-energy electrons are converted into heat energy due to collision of the electrons with the particles of workpiece material. The generated heat is used to melt the metal and create a weld pool to fill the weld area. Upon solidification, a quality weld bead is produced. The schematic diagram of the electron beam welding process is shown in Fig.\u00a07.16. The advantages, limitations and areas of applications are presented in the next section. 7.3 Electron Beam Welding 156 Some of the advantages of electron beam welding are as follows: \u2022 Thick plates can be welded in a single pass. \u2022 Generation of low distortion. \u2022 Less welding defect due to low weld contamination in the vacuum. \u2022 Narrow weld zone. \u2022 Narrow heat-affected zone. \u2022 It uses no filler metal. \u2022 High welding speed. \u2022 Welding of dissimilar metal can be achieved. \u2022 It produces high-quality weld and with high precision" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000012_6.2008-4509-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000012_6.2008-4509-Figure8-1.png", + "caption": "Figure 8-Labyrinth Seal Configuration (Four Tooth Design)", + "texts": [], + "surrounding_texts": [ + "American Institute of Aeronautics and Astronautics\n092407\n6", + "American Institute of Aeronautics and Astronautics\n092407\n7", + "American Institute of Aeronautics and Astronautics\n092407\n8\nTable 1 shows the seal combinations tested at Rexnord Corp. Figures 31 thru 36 show the performance of several seal combinations. Testing was completed at Rexnord which resulted in the seals been tested to 12,000 rpm. The approximate seal inlet temperature was 80o F" + ] + }, + { + "image_filename": "designv11_92_0001806_isiss.2018.8358114-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001806_isiss.2018.8358114-Figure1-1.png", + "caption": "Fig. 1: iXblue EVO-30L three-axis motion simulator structure.", + "texts": [ + " The purpose of this paper is to recall briefly the structure of a motion simulator (from a dynamic modelling point of view), and to introduce an innovative control scheme that attenuates these Coriolis and centrifugal effects. The presented experimental results demonstrate the efficiency of this algorithm. Furthermore the robustness of the associated control law make it perfectly suited to an industrial application. II. SYSTEM MODELLING A motion simulator generally has a gimbal-like structure, as shown in Fig. 1, which presents a three-axis equipment. 978-1-5386-0895-1/18/$31.00 \u00a92018 IEEE Such a machine, like most of all robotics systems can be modelled by means of Euler-Lagrange equations. Let us denote with: : axis angular positions vector : axis angular rates vector \u0393: torque vector (produced by actuators). The fundamental equation of robotics derived from EulerLagrange equation is [4]: + , + + = \u0393 (1) In (1), (inertial effects), , (Coriolis and centrifugal effects) are square matrices whose size is equal to the axes numbers", + " 3, but no estimation is done, since the compensation block reuses the parameters values that have been determined during the estimation procedure, and stored in a controller memory. Fig. 4 shows the control structure in the functional phase, which reveals to be very robust. IV. EXPERIMENTAL RESULTS These coupling disturbances from one or several axes to another axis, due to non-linear effects are abundant on a multi-axis motion simulator. This section addresses the evaluation of a specific coupling that occurs on a three-axis rotating table, as shown in Fig. 1. In some situations, it can reveal quite critical regarding the system stability performances: It includes the Coriolis effect on the inner axis due to the simultaneous middle and outer axis high velocity rate motion. Denote with , , the angular position and rate of the inner axis, , and , the angular position and rate of respectively the middle and the outer axis. The said Coriolis torque disturbance \u0393 affecting the inner axis can be well predicted by using a mathematical model derived from (1), and it has the following expression: \u0393 \u221d sin (2) Consequently, this torque is proportional to sin , implying that this Coriolis effect is a periodic disturbance at the frequency of the middle axis, modulated by the cross product of middle and outer axis velocities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000983_icelmach.2008.4800111-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000983_icelmach.2008.4800111-Figure5-1.png", + "caption": "Fig. 5. Vector diagram of PMSG showing the magnetomotive force vectors Ff, Fs, F\u03b4, flux vector \u03a6 and phase angles \u03b2 and \u03b4i.", + "texts": [ + " The main idea of loading method adopted for the calculation of load characteristics of stand-alone PMSG is exactly the same as in the case of loading method used for the prediction of electromechanical characteristics of PMSM. It is based on the possibility of calculation (after previous FEA of field distribution inside the machine) of an internal voltage Ei and internal power angle \u03b4i values, corresponding to the actual field distribution at each particular machine operating point, assuming that a modulus I1 and a phase angle \u03b3 of an armature current phasor I1 are known \u2013 see Fig. 3 and Fig. 5. A model of PMSG for FEA is prepared so that the A-phase axis of stator winding is exactly aligned with the d-axis of rotor \u2013 see Fig. 4. Additionally, we assume that the FEA model corresponds always to an instant of time t = 0, and that the instantaneous values of phase currents, which are the input variables for FEA, are described by equations: In such a case, appearing in (2) the phase angle \u03b2 of the stator current phasor I1, calculated with respect to the rotor d-axis: \u03b2 = \u03c0/2 \u2013 \u03b3, corresponds simultaneously to the space angle between a stator magnetomotive force (MMF) vector Fs and a rotor MMF vector Ff (Fig. 5). Thanks to that the exact positioning of stator MMF in relation to d-axis of rotor in the FEA model is very easy. Because of electromagnetic symmetry, the model of PMSG for FEA can be compressed to just one pole pitch. After computation of the magnetic field distribution in the generator for a given current modulus I1 and phase angle \u03b3, next the spatial distribution of magnetic vector potential along the air-gap (around the inner surface of stator) is calculated. This vector potential distribution corresponds to a resultant airgap flux vector \u03a6, which is a result of interaction between the rotor MMF Ff and the stator MMF Fs at this particular operating point of the generator \u2013 see Fig. 5. Calculated spatial distribution of magnetic vector potential along the air-gap is subjected to a harmonic analysis from which the Fourier coefficients a1 and b1 are obtained for the fundamental harmonic of vector potential. These coefficients represent respectively a half of d-axis flux and a half of q-axis flux in the generator air-gap, per unit depth [1] \u00f7 [3]. Knowing the Fourier coefficients a1 and b1 we can calculate the fundamental magnetic flux in the air-gap \u03a6, corresponding to given modulus I1 and given phase angle \u03b3 of the current phasor I1: 2 1 2 12 baL Fe +\u22c5\u22c5= , (3) where LFe is an axial length of motor core. Next, the modulus Ei of the internal voltage phasor Ei and the internal power angle \u03b4i \u2013 the phase angle of the resultant air-gap flux vector (see Fig. 5) are calculated using (4) and (5): sui kkzfE \u22c5\u22c5\u22c5\u22c5\u22c5= 11144.4 \u03a6 , (4) ( )11 abarctgi =\u03b4 . (5) In equation (4) f1 is the frequency of stator current, z1 is the number of winding turns in series per phase, ku1 and ks are the winding distribution factor and the slot skew factor respectively. Immediately after calculation of Ei and \u03b4i we can calculate the below listed quantities using equations (6) \u00f7 (15) delivered mainly from the phasor diagram shown in Fig. 3: \u2022 the modulus V1 of the phase voltage phasor V1 (and then the modulus VLL of line to line voltage at machine terminals): 111 RXjV \u22c5\u2212\u22c5\u2212= 11 IIEi , (6) \u2022 phase angle arg_V1 of the voltage phasor V1: ( )111 arg RXjarg_V \u22c5\u2212\u22c5\u2212= 11 IIEi , (7) \u2022 power angle \u03b4: 12/ arg_V\u2212= \u03c0\u03b4 , (8) \u2022 electromagnetic (internal) power Pei: ( )iiiei IIEP \u03b4\u03b2\u03b4\u03b2 sincoscossin3 11 \u22c5\u22c5\u2212\u22c5\u22c5\u22c5\u22c5= , (9) \u2022 electromagnetic torque Te: ( )nPT eie \u22c5\u22c5= \u03c030 , (10) \u2022 mechanical power at the generator shaft P1: mei PPP \u2206+=1 , (11) \u2022 active electrical power delivered by the generator P2: \u03d5cos3 12 \u22c5\u22c5\u22c5= IVP LL , (12) \u2022 reactive electrical power delivered by the generator Q: \u03d5sin3 1 \u22c5\u22c5\u22c5= IVQ LL , (13) \u2022 apparent electrical power delivered by the generator S: 13 IVS LL \u22c5\u22c5= , (14) \u2022 efficiency \u03b7: 12 PP=\u03b7 ", + " With given values of \u03b3 and cos\u03d5, all of the equations describing the stand-alone operation of PMSG with R or R-L type load are satisfied simultaneously by just one value of the armature current modulus I1. The correct value of modulus I1 is not known in advance and it must be found using iterative estimation methods. In the developed algorithm these iterative calculations are realized with the secant method. The iterative solution \u2013 correct value of modulus I1 \u2013 is found if the value of phase angle of the voltage phasor V1 calculated using (7) is equal to value calculated as the sum of angles \u03b2 and \u03d5 (see Fig. 3 and Fig. 5). For each successive step of iterative calculations for finding appropriate value of I1 one nonlinear FEA of magnetic field distribution inside the generator is carried out. Other principal assumptions taken into account in the developed algorithm (and software) for the calculation of load characteristics of stand-alone PMSG with R or R-L type load can be summarized as follows: \u2022 to derive the load characteristics for a range of output powers (e.g. VLL = f(P2)), calculations are realized for a given range of phase angles \u03b3 of the armature current phasor I1; \u2022 the load power factor cos\u03d5 is given in advance and it is constant in the entire range of considered generator loads; \u2022 for each particular value of armature current phase angle \u03b3 there is only one value of armature current modulus I1 that satisfies simultaneously all the generator equations derived from the phasor diagram shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001891_irec.2018.8362457-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001891_irec.2018.8362457-Figure2-1.png", + "caption": "Fig. 2. Two-dimensional full model.", + "texts": [ + " In practice, the Preisach density must be indentified. For this end, some experimental data are used. The static Preisach model is indentified using the material technique data by minimizing the quadratic error between measured data and simulated ones. It is shown in Fig. 1 that the calculated values of hysteresis cycle are quite similar to the experimental data. The obtained results show that the determined parameters are satisfactory. III. Application to a surface mounted permanent magnet motor The studied topology which is shown is Fig. 2 is an outer rotor surface-mounted synchronous permanent magnet motor with a radial air gap flux and distributed windings. This machine is a 3-phase, 4-poles, 48 stator slots configuration. The stator and rotor parts are made up of laminated iron core. In practice, a non-linear B-H curve of the laminated iron (M800-65A) is used to characterize ferromagnetic regions. The used permanent magnets are rare earth Neodymium-IronBoron. Linear material properties are used to characterize permanent magnet parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003990_j.enganabound.2019.04.008-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003990_j.enganabound.2019.04.008-Figure12-1.png", + "caption": "Fig. 12. Discrete models.", + "texts": [ + " (3) The radial clearance of the four-row cylindrical roller bearing is composed with the radial deformation of plate unit, the roughness amplitude of the race and the oil film thickness. (4) The contact angle of the single row tapered bearing is 49\u00b0, and the roller half-cone angle is 4\u00b0. (5) The mill roll is simplified to a cantilever beam, and the fixed end is the middle of the mill roll body. The reaction force is applied in the middle of the shaft block bottom, where is between the third row bearing and the fourth row bearing. The axially fixed constraint is applied at the corresponding position of the axial tailgate on the shaft block. Discrete models are shown in Fig. 12 . The numbers of elements and odes of shaft block both are 1240, and the numbers of mill roll are 1664 nd 1666, respectively. The numbers of contact elements and contact odes of shaft block both are 240, and also the mill roll. .3.2. Calculation results When the rolling force is applied as 100 kN, three situations of no ubrication, smooth condition and the roughness amplitude as 0.05 h 0 re simulated by the Bearing BEM under the EHL condition. The distri- utions of the pressure and load are obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003011_012067-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003011_012067-Figure6-1.png", + "caption": "Figure 6. Setup of the 3 point bending test.", + "texts": [ + " (UD) carbonfiber layers were used to increase the stiffness of the arm using longitudinal direction during the vacuum lamination process. Heat treatment was used to increase the adhesive and cohesion force between the connected parts. The final profile of the cross sectioned arm can be seen in figure 5. 5 1234567890\u2018\u2019\u201c\u201d For the bending test a smaller part was cutted out from the wing arm with a length of 400 mm. The bending test was made with Instrion 4482 general tensile machine using 3 point bending test. The figure 6 and 7 shows the setup of the bending procedure. The bending test was carried out until the specimen did not slip down on the support elements. The experiment shows that the chosen glues were stronger than the aluminium foam tensile strength even at the smallest cross-section of the specimens. Type 2 sectioned specimen break method can be seen in figure 8, the breaking plane is much above from the connected surfaces. As a result, both of the yield strengths and the tensile strengths were the Al 6061 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003675_022092-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003675_022092-Figure2-1.png", + "caption": "Figure 2 shows clearly the shape similar to rolling element interface on the Eulerian mesh surface. This deliberate meshing has two purposes: to make Eulerian mesh and roller shape adapt to each other and to render the initial material designation more convenient. The final nodes totaled 33560, units 23704, and Euler units (EC3D8R) 9072, of which 14632 are first-order hexahedral units (C3D8R).", + "texts": [ + " When the load is stable, oil film does not change with time, i.e. 0/ =\u2202\u2202 th , the equation at this point is: x hu x p u h x h \u2202 \u2202= \u2202 \u2202 \u2202 \u2202 6 3 (1.7) 3.Establishing Bearing Fluid-solid Coupling Model The three-dimensional model of the rolling bearing model is built using por/e. ABAQUS provides a variety of ways for meshing. What this paper applies is hexahedral structured meshing, which can handle more complicated models with more contact such as rolling bearings. An 8-node hexahedral de-integration unit (C3D8R) is used for meshing, as shown in Figure 2 [7-8]. CISAT 2018 IOP Conf. Series: Journal of Physics: Conf. Series 1168 (2019) 022092 IOP Publishing doi:10.1088/1742-6596/1168/2/022092 The outer ring of the bearing is fixed on the bearing housing so it does not rotate with the inner ring. For clearer observation of the boundary condition setting process, the Euler part in the model is hidden. Since the lubricating oil represented by the Euler body is difficult to flow out of the space between the outer wall of the inner ring and the inner wall of the outer ring due to the limitation of the various parts of bearing, the flow velocity of the Euler body on the two annular sides thereof should be limited" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001638_s1350-4789(18)30116-8-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001638_s1350-4789(18)30116-8-Figure5-1.png", + "caption": "Figure 5. Schematic of a mating ring with textured side wall and a prototype with thermocouple wires for laboratory test (Xia and Khonsari, 2013 and 2014).", + "texts": [ + " Within our group at the CeRoM, we have concentrated on modelling aspects associated with cavitation effects and the examination of lubrication regimes (Qiu & Khonsari, 2009; 2011a, b & c; and 2012) and optimisation of load-carrying capacity (Fesanghary & Khonsari, 2011, 2012; and 2013; Shen & Khonsari, 2013). All of these studies concentrated on texturing the lubricated seal faces. This section examines a different application of surface texturing with the primary objective of improving heat transfer performance. For this purpose, Khonsari and Xiao (2012) designed a stationary ring with dimples fabricated onto its sidewall, as illustrated in Figure 5, along with a prototype. This design has several rows of micro-dimples \u2013 etched by a laser \u2013 around the circumference of the mating ring. Six thermocouples were installed at different locations to monitor the temperature in real time. A series of laboratory tests were performed at the CeRoM with a seal testing apparatus to investigate the performance of the textured seal. Both conventional (plain mating ring) and dimpled surface rings were tested under identical operating conditions for comparison purposes (Xia and Khonsari, 2013)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002171_978-3-319-49574-3_12-Figure12.53-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002171_978-3-319-49574-3_12-Figure12.53-1.png", + "caption": "Fig. 12.53 Line trap components", + "texts": [ + " \u2022 Mounted on top of OHL circuit capacitor VTs. Mounted inside the line trap are a number of components the main item being the tuning device. This is for single frequency, double frequency, or wideband tuning and is selected for the range of frequencies to be communicated over the network. The protective device is a surge arrester connected across the coil terminals to protect the line trap against transient overvoltages. To prevent birds entering or nesting in the line trap, bird barriers can be fitted as shown in Fig 12.53. They are designed to ensure that the line trap has adequate cooling. Neutral earthing resistors (NERs), also called neutral grounding resistors, are employed in medium-voltage a.c. distribution networks to limit the current that would flow through the neutral point of a transformer or generator in the event of an earth fault. Earthing resistors limit fault currents to a value that does not cause any further damage to switchgear, generators, or transformers beyond what has already been caused by the fault itself" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002497_978-3-319-99620-2_12-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002497_978-3-319-99620-2_12-Figure5-1.png", + "caption": "Fig. 5. Parameters optimized using Response Surface Optimization method", + "texts": [ + " Researches should be aware that there will be a series of feasible and infeasible design schemes throughout the process of optimization and that they must choose the best scheme and test the rationality of its results. In a case of pressure vessel with two nozzles, four parameters \u2013 supposed to be the most influential on strain distribution under service load \u2013 were chosen: wall thickness of nozzle 1, wall thickness of nozzle 2, thickness of the vessel head and thickness of the cylindrical surface (Fig. 5). Upper and lower boundaries for each parameter have been selected after parameter choice. At the end of optimization process, the best parameter value should have been within defined boundaries. Next step in optimization, after parameters selection, was Design of Experiment (DOE). The purpose of a DOE was to gather representative set of data to compute RS and then run RSO. A set of design points was defined using different combinations of values for all four parameters within predefined boundaries" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002020_j.crme.2018.04.016-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002020_j.crme.2018.04.016-Figure4-1.png", + "caption": "Fig. 4. Impossibility to realize with a 3 d.o.f. robot model the sequence of three elementary joint movements imposed by MacConaill\u2019s version of Codman\u2019s paradox. (a) First abduction movement changing the initial axis sequence (1, 2, 3) into (1, 2\u2032, 3\u2032). (b) Second swing movement changing the axis sequence (1, 2\u2032, 3\u2032) into (1, 2\u2032, 3\u2032\u2032). (c) Resulting singular configuration due to the confusion between axis 1 and axis 3\u2032\u2032 making impossible the final extension back movement.", + "texts": [ + " MacConaill [11] reconsiders Codman\u2019s paradox by substituting to the two-step closed-loop movement a three-step closed-loop movement defined, for example, as follows: from same neutral initial position as considered earlier, a first abduction movement of \u03c0/2 is performed, then a second swing or horizontal flexion movement of \u03c0/2 is considered before a final vertical flexion, to bring back the arm at its initial position. Let us see how this sequence of three joint movements can be simulated with our robotic model: abduction is performed by a rotation around axis 1, then swing by a rotation around axis 2\u2032 but neither axis 2\u2032 or axis 3\u2032\u2032 is then able to realize the final back movement in extension, as illustrated in Fig. 4, due to the singularity configuration resulting from the confusion between axis 1 and axis 3\u2032\u2032 . The so-called gimballock phenomenon is well known in theory of mechanisms: the 3 d.o.f. ball-and-socket joint becomes singular when two of its axes come to be the same. But, in daily life experience, it is clear that no such gimbal-lock phenomenon occurs in human shoulder joint. The great originality of MacConaill\u2019s version of Codman\u2019s paradox would then be, according to us, to highlight the possibility for the shoulder complex to avoid this singularity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure10-1.png", + "caption": "FIGURE 10 overlay plot showing the relative positions of involute after radial and angular shifts. The involute after radial shift can be thought of to be a result of the angular shift", + "texts": [ + " The increase in the interface diameter causes an increase in the base circle diameter as well as the pitch diameter, causing radial displacement of the gear teeth. The radial increase of the base circle diameter and pitch circle diameter causes an increase in the base module as the number of teeth does not change, resulting in the change of base pitch. The increase in the base module manifests in positive profile slope. Profile deviation is measured normal to the tooth profile in the transverse plane. Figure 10 shows the profile deviation measurement. It shows the ideal un-assembled involute curve I, assembled involute curve III after a radial shift equal to the outward movement of the interface, and an unassembled involute curve II with an angular shift so that the assembled radially shifted curve intersects with it. The measured profile deviation, BC, due to the radial shift is measured along the normal to the ideal un-assembled curve and is the difference between the profile deviation along the normal due to the radial shift, AC, and the profile deviation along the normal due to the angular shift, AB" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002287_s00170-018-2422-y-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002287_s00170-018-2422-y-Figure4-1.png", + "caption": "Fig. 4 Analysis of minimum thickness when \u03b1f is greater than \u03b1 [24]", + "texts": [ + " Clearly, the workpiece thickness has a lower limit above which the postbending workpiece model can completely enclose the design model. Selection of workpiece material needs to be based on the lower limit of the thickness. Given the design parameters L1 (flange length of one side on the sheet metal workpiece), L2 (flange length of the other side on the sheet metal workpiece), and t (design model thickness), two intermediate process parameters are defined to describe the thickness increase amounts required for enclosing the design model along flanges L1 and L2, and they are denoted as t1 and t2, respectively (Fig. 4). Combining thickness increase parameters t1 and t2 with design model thickness t leads to the definitions of required workpiece thicknesses along flanges L1 and L2. The required workpiece thickness is the minimum thickness at each flange of the post-bending workpiece model to enclose the corresponding flange on the design model. At flange L1, the required workpiece thickness is defined as trequired1 (= t + t1), at flange L2 as trequired2 (= t + t2). It should be noted that the required workpiece thicknesses trequired1 and trequired2 are not necessarily equal as thickness increase amounts t1 and t2 depend on the angular deviation along each flange between postbending workpiecemodel and the designmodel", + " In the author\u2019s previous work [24], the calculation of optimal thickness to was developed when \u03b1f is greater than \u03b1, and was used to guide the selection of workpiece material for angular accuracy enhancement. In this subsection, the method in the author\u2019s previous work [24] for calculating optimal thickness to when \u03b1f is greater than \u03b1 is first recapitulated, and then it is extended to the case where \u03b1f is less than \u03b1. (1) \u03b1f is greater than \u03b1 In the author\u2019s previous work [24], the design model is fitted into the post-bending workpiece model, and the optimal thickness is analytically found to completely enclose the design model. As illustrated in Fig. 4, the bottom boundary of the designmodel is first made to touch the bottom boundary of both flanges of the post-bending material model. This step relates these two models by providing the baselines to the thickness increase amounts t1 and t2. Following that, reference lines are created parallel to the bottom boundary of both flanges of the post-bending material model while tangent to the inner radius of the design model. In such a way, the distance between the reference line and the bottom boundary of each flange establishes the thickness increase at each flange. Given actual thickness ts =Max (trequired1, trequired2) and required workpiece thickness trequired1 (=t + t1), trequired2 (=t + t2), the actual thickness is further written as ts = t +max (t1, t2), where t is the thickness of the design model. With the definition of minimum enclosing thickness as to =Min [Max (trequired1, trequired2)], the objective of minimizing thickness min{t + max(t1, t2)} can be further written as Minimum thickness \u00bc t \u00femin max t1; t2\u00f0 \u00def g From Fig. 4, the relation of t1, t2 with other parameters is found t1 \u00bc L1 sin \u03b81 \u00fe 2Ri sin 2\u03b81 2 ; \u00f03\u00de t2 \u00bc L2 sin \u03b82 \u00fe 2Ri sin 2\u03b82 2 ; \u00f04\u00de \u03b81 \u00fe \u03b82 \u00bc \u03b1 f \u2212 \u03b1 \u00bc \u0394\u03b1 \u00f05\u00de When t1= t2, the thickness is minimized.With the relations above, the following equation leads to the solution of \u019f1 and \u019f2, which describe the orientation of design model within the post-bending workpiece model. L1sin\u03b81 \u00fe 2Risin \u03b81 2 sin \u03b81 2 \u00bc L2sin\u03b82 \u00fe 2Risin \u03b82 2 sin \u03b82 2 Let parameters \u03a6 be defined as sin\u22121 Riffiffiffiffiffiffiffiffiffiffiffiffi L12\u00feRi 2 p , \u03b8 as sin\u22121 Riffiffiffiffiffiffiffiffiffiffiffiffi L22\u00feRi 2 p , T1 as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L12 \u00fe Ri 2 p , and T2 as ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi L22 \u00fe Ri 2 p , the following equation is obtained: With parameter \u03b2 defined as \u03b2 \u00bc tan\u22121 T2sin \u03d5\u00fe\u03b8\u2212\u0394\u03b1\u00f0 \u00de T1\u00feT2cos \u03d5\u00fe\u03b8\u2212\u0394\u03b1\u00f0 \u00de \u00fe 0 if T 1 \u00fe T2cos \u03d5\u00fe \u03b8\u2212\u0394\u03b1\u00f0 \u00de\u22650 \u03c0 if T1 \u00fe T2cos \u03d5\u00fe \u03b8\u2212\u0394\u03b1\u00f0 \u00de<0 ; angle\u019f1 can be calculated as \u03b81 = \u03d5 \u2212 \u03b2", + " When the actual workpiece thickness T deviates from the optimal thickness to, angles \u019f1 and \u019f2 are not fixed; instead, each of them belongs to a feasible set. Let \u019fj (j = 1, 2) collectively represent angles \u019f1 and \u019f2. The range of angle \u019fj is denoted by \u019fj \u03f5 [\u019fjmin, \u019fjmax]. Let \u2202Wa and \u2202Wd denote the boundaries of the actual workpiece and the design model, respectively. The problem of finding the range for \u019fj is then formulated as the determination of its limits subject to two constraints such that: \u019f j \u03f5 \u03b8 j;min; \u03b8 j;max Subject to \u2202Wd \u2201 \u2202Wa \u2211 2 j\u00bc1 \u03b8 j \u00bc \u0394\u03b1 \u00f011\u00de For the first case discussed in this section where \u03b1f is greater than\u03b1 (Fig. 4), rewrite the expression of thickness increases t1 and t2 collectively into tj. t j \u00bc Ljsin\u03b8 j \u00fe 2Risin 2\u03b8 j 2 , where j = 1, 2. With the definitions of T j \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Lj 2 \u00fe Ri 2 q and \u03c6\u2217 = atan2 (\u2212Ri, Lj), the ex- pression changes to t j\u2212Ri \u00bc T jsin \u03b8 j \u00fe \u03c6* Hence, \u03b8 j \u00bc sin\u22121 t j\u2212Ri T j \u2212\u03c6* \u00f012\u00de Let \u0394t be the difference between the optimal thickness to and the selected workpiece thickness T such that \u0394t = T \u2212 to. The first constraint \u2202Wd \u2201 \u2202Wa leads to the maximum \u019fj, such that \u03b8 j \u00bc sin\u22121 \u0394t\u2212Ri T j \u2212\u03c6*, whereby the boundary of design model barely touches the boundary of the selected workpiece", + " Therefore, the objective is to minimize the difference of deviation angles \u019f1 and \u019f2; hence, it is written as Min abs (\u019f1 \u2212\u019f2) Subject to \u019f1 \u03f5 [\u019f1min, \u019f1max] \u019f2 \u03f5 [\u019f2min, \u019f2max] \u019f1 +\u019f2 =\u0394\u03b1 Workpiece length can be estimated based on the way of defining minimum enclosing thickness. The estimation uses a pure geometric approach for simplification purposes. Phenomena such as thinning and bulging at the deformation zone are neglected. Depending on the relation of final bend angle \u03b1f and design bend angle \u03b1, workpiece length can be estimated as the summation of both flange lengths and arc length at the deformation zone. When \u03b1f is greater than \u03b1 as illustrated in Fig. 4, the workpiece length can be calculated using the following formula. L1 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b81 2 cos\u03b81 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b81 2 \u00fe L2 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b82 2 cos\u03b82 \u00fe Ri \u00fe t\u00f0 \u00detan \u03b82 2 \u00fe \u03c0\u2212\u03b1 f Ri \u00fe t\u2212 top 2 When\u03b1f is less than\u03b1 as illustrated in Fig. 5, the workpiece length can be calculated in a similar way with the following formula L1\u2212Ritan \u03b81 2 cos\u03b81 \u2212 Ritan \u03b81 2 \u00fe L2\u2212Ritan \u03b82 2 cos\u03b82 \u2212 Ritan \u03b82 2 \u00fe \u03c0\u2212\u03b1 f Ri \u00fe top 2 Given the nature of machining process used in this research, the process planning requires a compensation step before toolpath planning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure44-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure44-1.png", + "caption": "Fig. 44 Input-output characteristics", + "texts": [ + " This device is designed to make it possible to measure torque quite precisely by virtue of a low friction air bearing and torque loading with tension force. A commercially available high-precision torque meter is also embedded for cross-checking. We measured torque-sensor output while increasing input torque gradually to the maximum and then decreasing to the maximum of opposite direction. It was confirmed that the measurement variation was less than 0.1% after 5 cycles of the increasing-and-decreasing measurement procedure. Figure 44 shows Type B\u2019s characteristic relationship between signal output and load torque. Nonlinearity is about 0.8 (%/FS). Fig. 42 Subassembly Fig. 43 Torque sensor evaluation device Fig. 46 Torque-servo module We also investigated to what extent each sensor resistance varies along with temperature variation by means of a commercially available constant temperature bath. As a result of the experiment, it was confirmed that TCR (temperature coefficient of resistivity) was less than 100 (ppm/\u0131C (0-85\u0131C)) (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002701_s1068798x18090125-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002701_s1068798x18090125-Figure1-1.png", + "caption": "Fig. 1. Determining the influence of vibration on the roller skew.", + "texts": [ + " In vibrational and rotary motion of the joint as a result of nonuniform double-speed rotation of the Cardan shaft and resonant vibration (produced by excitation sources such as the motor, the primary transmission, the gear box, and the driving wheels), a variable torque is applied, resulting in vibration in individual components of the joint\u2014in particular, the needle. Under the action of this torque, the torque due to the frictional forces, which skews the needle, constantly changes sign [7, 14, 15]. However, since the variable torque due to the frictional forces acts on the needle in contact with the journal, it also performs elastic vibration relative to the axis of the crosspiece pins. Such vibration is facilitated by the radial and axial freedom of motion (within the gaps). The needle vibration is illustrated in Fig. 1. For the sake of simplicity, we assume that the mass of the needle is concentrated in the points where the forces act on both sides of the needle\u2019s center (Fig. 1). The resistance associated with needle vibration by an angle \u03c8 is assumed proportional to the angular velocity [7, 16]. In that case, the needle\u2019s equation of motion takes the form (1) where Jne is the needle\u2019s moment of inertia; \u03c8 is the needle\u2019s skew angle; \u03be is the damping coefficient; K is the drag; Mme is the mean constant component of the load torque on the cardan joints; Mv is the variable component of the load torque; \u03c9in is the frequency of the inducing force; and t is the time. The intrinsic vibrations of the needle due to the contact friction are disregarded, since they are negligible in comparison with the forced vibrations [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000881_pi-b-2.1962.0229-Figure13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000881_pi-b-2.1962.0229-Figure13-1.png", + "caption": "Fig. 13.\u2014V.H.F. notch aerial in Scimitar tail fin.", + "texts": [ + " Over the civil aircraft band, 118-132Mc/s, a simple coaxial feed was sufficient, but a short-circuited cable stub was added at the appropriate point along the coaxial cable to match the full band of 100-156 Mc/s. Impedance curves for the two bands are shown in Fig. 12. It should be emphasized that the Herald notch radiation patterns FEEDER ACROSS NOTCH DIELECTRIC COVER FIN LEADING EDGE One unexpected difficulty was met in the application of a notch aerial to the Scimitar aircraft, whose tail fin is shown in Fig. 13. The gap between the top of the fin and the rudder mass-balance acted as a parasitic notch, resonant within the required frequency band, and prevented the desired current flow around the periphery of the tip of the tail structure, so that 436 BURBERRY: PROGRESS IN AIRCRAFT AERIALS there was no rearward radiation as the radiation patterns of Fig. 14 indicate. This effect has been used to give directivity to a notch aerial in an application described in a later Section of the paper. Where structural limitations prevent the use of a straight notch, it is possible to obtain the required length by bending or folding the notch, as Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002171_978-3-319-49574-3_12-Figure12.16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002171_978-3-319-49574-3_12-Figure12.16-1.png", + "caption": "Fig. 12.16 Typical top-core current transformer", + "texts": [ + " Current transformers consist of five main components, in the top-core oil-insulated design: \u2022 The primary head housing the required cores or ring transformers which are assembled from core rings of iron on which is wound a precise number of wire turns \u2022 The primary conductor passing through the center of the head \u2022 A secondary terminal or junction box where the wire tails from the ends of the winding tails of the CT cores are terminated and where the interface to the protection and control cabling is located \u2022 The insulating housing which can be porcelain or polymeric and the base frame for mounting the CT on a support structure \u2022 Insulating oil which fills the void between the core, tails, head, and support insulation In Fig. 12.16, it can be seen that there are ring transformers mounted in the head of the instrument transformer. These have a specific transformation ratio, e.g., 2000:1, (2000 turns of wire around the magnetic core to 1 primary single turn conductor), so 2000 A flowing in the primary circuit would translate into 1 A flowing in the secondary output of the instrument transformer. This can then be used by protection and control equipment as required to monitor the safe operation of the substation equipment and associated network (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003348_icarm.2018.8610860-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003348_icarm.2018.8610860-Figure2-1.png", + "caption": "Fig. 2. The MoMaCo mobile manipulator configuration and the reference frames of the system", + "texts": [ + " \u2018semi\u2019 is caused that the deviation between the actual position of the building system components and the ideally planned position may occurs sometimes. Although the shape of the house is known, for a mobile manipulator moving in the room, the position and pose of end-effector will be affected by the uneven ground. III. The studied mobile manipulator in construction (MoMaCo) is mainly composed of a mobile platform and a manipulator arm mounted on a lifting column which is used in construction field for decoration engineering, as shown in Fig.2. The mobile platform comprises four driving Mecanum wheels so that it can move in all directions which make it has locomotivity even in a narrow environment [13]. A 7-DOF cooperative manipulator arm KUKA LBR iiwa whose base can 978-1-5386-7066-8/18/$31.00 \u00a92018 IEEE 662 move up and down along with lifting column is mounted onto the mounting plate. The position and the direction of the mobile platform can be obtained with the laser scanner sensor which is mounted on the front of mobile platform. Considering that the ground of the environment may not be flat, the posture of lifting column relative to the horizontal plane can be measured by an inclination sensor mounted on the column", + " For the lifting column, its pose relative to the horizontal plane will be affected by the uneven ground when the mobile manipulator moves on the room, as shown in Fig.3(b). When the lifting column inclines with an angle of degrees relative to the vertical line, the position of the end of the column will change by D=S*sin . In case of S=1m, =1\u00b0, the value is 17.45mm. It may be more larger to the tip of manipulator. It is very necessary to obtain the incline data of the lifting column even though the value is small. The kinematics analysis of studied mobile manipulator system consists of the following six frames, as shown in Fig.2. \u2022 Wall coordinate system (WCS) {W}: OWXWYWZW. It is fixed on the wall i.e. the environment. The ZW axis is along vertical direction. The XW axis is along intersection line between vertical wall and horizontal plane. \u2022 Mobile base coordinate system (BCS) {B}: OBXBYBZB. According to the approximation above, the plane OBXBYB is always considered as being horizontal even though the mobile platform moves on an uneven ground. The origin OB locates on the bottom of the lifting column. The ZW axis is along vertical direction and the YB is parallel with the front surface of the column, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000814_s11029-009-9095-4-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000814_s11029-009-9095-4-Figure6-1.png", + "caption": "Fig. 6. Vari a tion in tan q within the lim its of one wind ing coil of the branch at t = 5; b3 = 30\u00b0.", + "texts": [ + " (20) Let us con sider the sec ond re stric tion. For a shell of rev o lu tion, tan q x b + b = d dz k k sin cos2 1 2 2 , where x b= rsin . For a cir cu lar to rus, d dz d d R x x = \u00d7 J J1 1 1 sin , k k t1 2 1 1= + sin J . Tak ing into ac count re la tion (12) and per form ing some trans for ma tions, we have tan q l l l l = + + + + \u00e9 \u00eb\u00ea \u00f9 \u00fb\u00fa cos ( ) ( ) sin ( ) J J 1 2 2 1 2 1 2 2 1 1 1t t , where l b = + + \u00e6 \u00e8 \u00e7 \u00f6 \u00f8 \u00f7 tan 3 1 1 1 cos sin J J H t t . The vari a tion in tan q within the lim its of one coil is shown in Fig. 6. To en sure the equi lib rium of fil a ments, the re stric tion | tan frq|max \u00a3 k must be ful filled, where k fr is the fric tion co ef - fi cient. With a high de gree of ac cu racy, it can be as sumed that | tan q|max cor re sponds to sin J 1 0= . Ne glect ing ad di tion ally the value of l2 in com par i son with unity, we ar rive at the fol low ing estimate: tan frb3 2 \u00a3 +k t H( cos )J . (21) At k fr \u00bb 0.2-0.3, re stric tion (21) is stron ger than (20) and must be taken into ac count in de sign ing a branch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000288_1542-6580.2079-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000288_1542-6580.2079-Figure1-1.png", + "caption": "Figure 1. Schematic diagrams of electrochemical batch reactor: a) Divided Cell b) Undivided Cell", + "texts": [ + " All the experiments have been carried out at 25 \u00b1 1 oC. In all experiments, the electrochemical cell has been washed with nitric acid, distilled water and finally triple distilled water before recording the cyclic voltammogram. Electrochemical activation procedure has been adopted whenever the electrode reproducibility was lost. For calculating experimental data, care has been taken to detect the background current at each cycle. The electrochemical batch reactors have been used for direct reaction studies in divided (Figure 1a) and undivided (Figure 1b). The divided cell has been provided with two (anolyte and catholyte) compartments. The anolyte and catholyte compartments have been separated by Nafion 423 (cationic exchange membrane) which has a capacity of 100 mL and 500 mL respectively. The cell has been made up of PVC. The cell contents have been agitated by a glass stirrer connected to electrical motor provided with a speed control unit. Brought to you by | The University of Texas at Austin Authenticated | 128.83.63.20 Download Date | 9/14/13 5:57 AM For preparative studies, Lead/Lead dioxide (0.24 dm2) and titanium substrate insoluble anode, TSIA (0.24 dm2) with TOx-RuOx-IrOx coating (Subbiah, 1990), have been used as anodes and stainless steel (0.24 dm2) was the cathode. Other instruments such as DC regulated power supply (APLAB DC Power Supply 0-20A 0-30V), stirring unit, heating source, cryostat (Julabo Make F10 model) have been employed for preparative electrolysis. The cell set up used for direct studies has been shown in Figure 1. The preparative studies of the oxidation of dimethylsulphoxide to dimethylsulphone have been carried out in a divided rectangular cell as per the experimental conditions listed in the Table 1. For the oxidation of DMSO, aqueous MSA of concentration varying from 0.25 M \u2013 2.0 M have been used as electrolyte. DMSO has been dissolved in MSA to get the desired concentration which has been used as anolyte and aqueous MSA has been used as catholyte. Electrochemical oxidation studies were conducted to optimize various reaction parameters such as current density, acid concentration etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002304_urai.2018.8441777-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002304_urai.2018.8441777-Figure5-1.png", + "caption": "Figure 5. Simulation results of the blending trajectory. (Red, green and blue arrows indicate x, y, and z axes of the waypoints and end-effector)", + "texts": [ + " The computational results of Algorithm 2 under the above conditions are described in Table 2. By the time synchronization process, the maximum angular velocity and blending radius of the orientation trajectory were re-computed because the duration time of the orientation trajectory was smaller than that of the position trajectory (tpos > tori). As a result, the duration time of the two trajectories was synchronized (tpos = orit ). Using the above results, the blending trajectory was computed as shown in Fig. 5. The position trajectory of the end-effector was generated smoothly and continuously without passing through the intermediate waypoints P1 and P2, and so was the orientation trajectory. Through the above simulations, the results of each axis were obtained as shown in Fig. 6. In Fig. 6, the blue background represents the acceleration and deceleration section of the blending trajectory, and the green and red background denote the cruise velocity and blending sections, respectively. As shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000498_s1672-6529(09)00003-1-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000498_s1672-6529(09)00003-1-Figure1-1.png", + "caption": "Fig. 1 General arm model.", + "texts": [ + " They are complete and can be cast in any variational week form for numerical solution, including a generalized nonlinear finite element, or simpler beam element models. This convenience will be utilized in two-link model of the human arm that is presented in the next section. 3 Elastic two-link and hand load model of human arm We now derive the model used for a two-segment elastic model of the human arm with a non-articulated hand load. The generality of the model allows one to incorporate muscle effects as either loads attached through the tendon at points of origin and insertion or as loading torques at the shoulder and elbow. The model is depicted in Fig. 1. The domain of each beam is one dimensional. The independent coordinates are x11 and x21 measured from the root of beam B1 and B2, respectively, along the undeformed neutral axis of each beam. The \u201cspecial\u201d point of beam B1 is labeled p1 (the pivot) and p2 (the joint between the two beams) is for beam B2. The coordinate frames, denoted with B1 and B2, are attached as shown in Fig. 1. At the root of each beam (B1 and B2) there are massless hubs to which torques M1 and M2 are applied. The angular position of frames B1 and B2 are q1 and q5. Beam deflections are measured with 11 11 11 \u02c6( , )u x t b and 21 21 21 \u02c6( , )u x t b (elongation), and with 12 11 12 \u02c6( , )u x t b and 22 21 22 \u02c6( , )u x t b (flexure), as shown in Fig. 1. The beams have mass per unit length (xi1), total lengths are L1 and L2, cross sectional area A, area moment of inertia (xi1), and Young\u2019s modulus E. It is assumed that large deflections and rotational inertia are pertinent, but not shear deformation. Therefore, the beams are modeled with Rayleigh beam theory with elongation. The cross-sections of each beam are assumed symmetric about the neutral axis. The intrinsic mass moment of inertia of the beam cross-section will be taken as 3 3 \u02c6 \u02c6( ) oi I x i i , and 3\u0302i is normal to the plane of the problem", + " The results are presented for the case where the shoulder and elbow are controlled by torques with simple proportional control on joint angles with viscous damping. The model allows one to incorporate muscle loads transmitted through the tendon at points of origin and insertion, but we reserve those studies for other work[18]. The properties chosen for the arm are shown in Table 2. For this simulation, a solid circular cross-section was assumed as well as identical lengths for both links. The arm was initially hanging relaxed, straight down, see Fig. 1. The arm was then commanded to rotate the first link to horizontal and the second link to vertical in a curling-like motion. The simulation was run for 5 seconds duration. The responses for the regular generalized coordinates and the coordinates used to discretize the beams (see Eqs. (62) and (63)) will be reported. shown, respectively. The elastic response can be seen in the initial response before 3 seconds of elapsed time. The ringing from the proportional control dominates the remaining response" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001699_1077546318767559-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001699_1077546318767559-Figure2-1.png", + "caption": "Figure 2. Dynamic model of the geared rotor system.", + "texts": [ + " The results show that the acceleration of the box can reflect the meshing state of gear pairs well, which will be helpful in the vibration control and condition monitoring of the equipment. The structure of the geared rotor system in an integrally geared compressor is shown in Figure 1, which was obtained through the similarity design method. The geared rotor system is composed of five parallel shafts which are meshed by a helical gear including one input shaft, one intermediate gear shaft, and three high speed output shafts. Both ends of the five shafts are respectively supported by an oil film bearing. The dynamic model of the five shafts geared rotor system is shown in Figure 2. The gears on I, M, O1, O2, and O3 are numbered 1\u20135 with angular velocities !1 to !5, respectively, where time-varying mesh stiffness kij\u00f0t\u00de, backlash 2bij, and transmission error eij\u00f0t\u00de are included. kij\u00f0t\u00de and cij are the mesh stiffness and damping between gear i and j. eij\u00f0t\u00de is the transmission error between gear i and j. bij is the single side backlash between gears i and j. KiLand KiR are the bearing stiffness matrix of the left bearing and the right bearing of shaft i, while CiL and CiR are the support damping matrix of the left bearing and the right bearing, respectively. The geared rotor system shown in Figure 2 can be divided into different elements to be modeled and analyzed by the finite element method. The corresponding gear nodes are connected by gear elements, and every shaft is supported by two bearing elements. The bullgear and other additional parts of the counterweight plate fixed on the shaft are simplified into a lumped mass element. The dynamic model of the helical gear is shown in Figure 3, where i and j are driving and driven gears, respectively, while !i and !j also can be expressed by _ xi and _ xj" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002242_978-3-319-99262-4_29-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002242_978-3-319-99262-4_29-Figure4-1.png", + "caption": "Fig. 4. Seals test rig", + "texts": [ + " The solution from one field serves as the initial value for the other. Finally, Eq. (6) was fitted to the simulated values of the forces to calculate the rotordynamic coefficients. This section describes the experimental setup for analyzing rotor seal systems at the Chair of Applied Mechanics at the Technical University of Munich, as in [21]. First, the seals test rig is presented, then the dynamic behavior is analyzed and compared to the simulation results. The experimental analysis is examined on the seals test rig (see Fig. 4). The main components are a flexible shaft and a mass disk (1) with two symmetricallyarranged liquid annular seals (2) in the middle (see details in Fig. 5). Eddy current sensors for measuring the displacement (6) and a piezo force platform (7) are arranged in the seals stator housing (8). The fluid is injected between the two seals with a maximum pressure of 100 bar. The rotor runs at over-critical speed above the first (sealless) natural frequency \u03c91 up to 12,000 rpm. An active magnetic bearing (3) is used as an exciter (2D shaker) at the shaft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001626_j.measurement.2018.03.009-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001626_j.measurement.2018.03.009-Figure11-1.png", + "caption": "Fig. 11. Orbit plot for delevitation 3 @ 4589 r/min (rocking motion with small bounces) [17].", + "texts": [ + " Once again, note the presence of noise in the experimental results. The comparison between simulated and experimental results displays very similar behaviour for the experimental and simulated RDEs, with the exception of the amplitude of the rocking motion. The difference in amplitude of the rocking motion is attributed to the fact that the bearing and bearing housing stiffness and damping values in the simulation model do not exactly match the real world values. An RDE with an initial rotational speed of 4589 r/min is shown in Fig. 11. The experimental and simulated results seem similar with the exception of the small bounces on the left side of the orbit plot of the simulated results. This is again attributed to the fact that the stiffness and damping characteristics of the simulated model do not exactly match those of the experimental setup. Both the simulation and the experimental results show a rocking motion. The results of an RDE with an initial rotational speed of 5097 r/min are shown in Fig. 12. Comparing the rotor behaviour clearly shows that both the simulation and the experimental results predict a tendency to go into a forward whirling motion, but both only experience a period of bouncing, and then settle to a rocking motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001471_1.4038952-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001471_1.4038952-Figure1-1.png", + "caption": "Fig. 1 Solid and hollow rollers under compression", + "texts": [ + "4038952] Keywords: hollow roller bearing, contact deformation, hollowness ratio, elastic approach Introduction The line contact deformation of a cylinder compressed by two plates is a basic building block in roller bearing deformation equations, and it is the foundation of the bearing load distribution calculation. The total approach of the two plates is determined by the sum of deformation in the roller and the two plates. Many equations have been proposed by researchers [1\u201312]. A diagram of solid and hollow rollers compressed by two plates is shown in Fig. 1. When a solid roller is compressed by two plates, the elastic approach is equal to the difference of displacement between points 4 and 1, which can be calculated by the following equation [5\u20139,12]: dr \u00bc kq ln 4r kq 1 (1) where k \u00bc 2\u00f01 l2\u00de=pE; l and E are the Poisson\u2019s ratio and Young\u2019s modulus; r is the outer radius of roller; q is the line load. However, for hollow rollers, the calculation of deformation is more complicated. The parameter of hollowness ratio (hr \u00bc ri=r), which is first proposed by Bowen and Bhateja [13], is the ratio of the bore radius (ri) to the outer radius (r) of hollow roller. The elastic approach of hollow roller (dhr) includes two parts: the contact deformation (dc) and the bending deformation (db). In Fig. 1, the difference of displacement between points 4 and 3 equals the difference of displacement between points 2 and 1, which are both equal to dc=2; the bending deformation db is equal to the difference of displacement between points 3 and 2. Currently, most of the formulas for computation of hollow roller\u2019s deformation are simply the sum of contact deformation and bending deformation. But in Refs. [13\u201315], formula (1) is directly adopted to calculate the contact deformation of hollow roller without considering the influence of roller\u2019s hollowness ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000102_s1052618808020040-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000102_s1052618808020040-Figure3-1.png", + "caption": "Fig. 3.", + "texts": [ + " The process is further controlled by truncated equations (3) and ends at stable sta- y\u0307 b11 \u2013b 2sf \u03c0 1\u2013 \u03c90 3\u2013 YV 2 A* 3\u2013 , Y\u2013 1 V /\u03c90A*( )2 \u2013[ ] 0,5\u2013 \u03b1, b12cos e \u03d5*,cos= = = b13 = 2sf \u03c0 1\u2013 \u03c90 3\u2013 VYA* 2\u2013 , b21 = \u2013s 2\u03c90( ) 1\u2013 A* 2\u2013 \u03d5*, b22cos = \u2013e 2\u03c90( ) 1\u2013 A* 1\u2013 \u03d5*,sin b31 2gf \u03c0 1\u2013 \u03c90 1\u2013 VYA* 2\u2013 , b33 \u20132gf \u03c0 1\u2013 \u03c90 1\u2013 YA* 1\u2013 , b23 0, b32 0.= = = = B1 0, B3 0, B1B2 B3\u2013 0, B1> > > \u2013 b11 b22 b33+ +( ),= B2 b11b22 b11b33 b22b33 b23b32\u2013 b12b21\u2013 b13b31,\u2013+ += B3 b11b23b32 b12b21b33 b13b22b31 b11b22b33\u2013 b12b23b31\u2013 b13b21b32.\u2013+ += x\u0307 y\u0307 JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 37 No. 2 2008 EFFECT OF A VIBRATION DRIVE ON VIBROTRANSPORTATION 121 tionary point D. This process has the form shown in Fig. 3a, where A1 and A2 are the vibration amplitudes at points C and D, respectively. It should be noted that, although at each stage the motion is described by different equations with a different number of degrees of freedom (a system with a variable structure), as a whole, we obtain a solution of the Cauchy problem with a known accuracy. In the case under consideration, the equations for systems with a variable mass are not used since the relative velocity of masses M and m at the moment of separation is zero", + " One can conclude that, for vibration amplitudes that are smaller than boundary (5), the regime of the joint motion of the mass and the base is the sole stationary regime. Returning to the transient process, after point E, the vibration amplitude decreases according to (3) and tends to point F until condition (5) is violated. After this, the system loses one degree of freedom and transforms into a linear oscillator that tends again to the unrealizable stationary regime (the broken line on curve 2). This process is shown schematically in Fig. 3b, where A3 is the amplitude at point E. Similar regimes exist for frequencies larger than those at point B (the broken line on curve 1 in Fig. 2a). For description of these regimes, coordinates rather than their envelopes (3) must be known, and, therefore, it is not possible within the framework of this study. Nevertheless, it is proven that vibrotransportation is possible within broader ranges of amplitudes and frequencies than in the classical theory. The described regimes are characterized by \u201csoft\u201d excitation of vibrations, i", + " The exact solution shows that, above and close to boundary (5), only regimes with finite-length stops in relative motion are possible. The boundary of existence of regimes with only instantaneous stops may be easily found from the formulas reported in [1] or from the diagram presented in [7] (8) where k = const is found using two parameters f and \u03b1 in [1]. According to (8), curve 4 (Fig. 2) always lies above curve 2 (shown by a dot-dash line). These stops are possible in segments AH and IB (Fig. 2a) of resonance curve 1 between curves 2 and 3. A \u201csoft\u201d regime is realized, which is similar to that in Fig. 3b, with vibrotransportation and vibration amplitudes located between resonance curve 1 and boundary 2 (thin arrow). If excitation amplitude E0 decreases, a situation is possible in the system shown in Fig. 2b where curve 2 passes above oscillator amplitudes (curve 3). The onset of regimes with vibrotransportation is possible only on the AB segment of curve 1 through \u201chard\u201d excitation of vibrations. The initial conditions for this regime are set by the corresponding selection of A(0), (0), \u03d5(0) \u2192 \u03c8(0) and inverse transform according to (2) to initial coordinates in (1). On segment HI of curve 1, stable vibrotransportation regimes are implemented, and on segments AH and BI, regimes of the type shown in Fig. 3b are implemented. \u201cIsolated\u201d segments AB of resonance curve 1 may also be reached by \u201cextending\u201d the vibration regime by increasing and subsequently slowly decreasing the excitation amplitude. Resonance curve 3 of the linear oscillator may be implemented in full as in Fig. 2b; therefore, it is rather difficult to find the \u201chard\u201d vibrotransportation regimes in an experiment without calculations. If the excitation amplitude is decreased further, curve 1 goes down (Fig. 3b) and has only one common point J with known amplitude and frequency \u03c90 on curve HI. Therefore, the possible regulation of vibrotransportation by changing the voltage amplitude E0 is limited according to the first expression in (4). 4. We now consider specific features of the obtained regimes. If mass m is fairly small (m/M ~ \u03b5), the term with Coulomb friction in the first equation in (1) will be of the second order of smallness and, in the first approximation, the mass motion does not affect vibrations", + " In the general case, the classical theory may be used to adjust the vibrotransportation velocity for nonstop regimes [1]: (9) In the second formula of (4), \u03c90 may be replaced without loss of accuracy by \u03c9 since in the third expression in (3) this replacement holds with an accuracy of \u03b52 owing to the formulas \u03c90 = \u03c9 \u2013 \u03b5\u2206 and /\u03c90A) = /\u03c9A) + \u03b5\u2206(\u2026) + \u03b52\u22062(\u2026) + \u2026. Thus, the exact solution of (9) differs from approximate solution (4) by a value of the order of \u03b52. It should be noted that expressions (4) and (9) for system (1) have the same order of accuracy. Other regimes for which the vibrotransportation shown in Fig. 3b is possible may not be recommended for practical use due to their possible stochasticity although it frequently appears in different systems between unstable limiting cycles rather than between stable ones. The regimes with hard excitation ensure the lowest voltage and amperage on the vibration exciter winding. They may be recommended for continuous vibrotransportation. The regimes with soft excitation are also suitable for dosing operations since they may be implemented when a mechanism is switched on" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002254_isie.2018.8433668-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002254_isie.2018.8433668-Figure8-1.png", + "caption": "Fig. 8. B4 Fed TPSIM.", + "texts": [ + " The electromagnetic torque developed in the rotor is shown in Fig. 5. and rotor speed is determined using the mechanical modelling concept of motor shown in Fig. 6. [14]. III. FOUR SWITCH TWO LEG INVERTER FED IM DRIVE The four-switch two leg also known B4 inverter can only have four topologies. These four topologies are appeared underneath in Fig. 7. All these distinct four topologies produces non-zero voltage also known as active states. The two phase symmetrical induction motor drive connections with a B4 two phase voltage source inverter is shown in Fig. 8. [15]. The intrinsic active voltage vectors produced by the two-leg inverter coupled with TPSIM are shown in space vector (SV) diagram Fig. 9 while Table. I summarizes switching states with their space vector magnitude and are clearly quadrature displaced. IV. DTC FOR B4 FED TPSIM DRIVE Different DTC techniques have been studied in the literature for single phase induction motor drives [16], [9] but not a lot of work has been done in case of TPSIM. Table II shows the vector selection table for conventional DTC strategy based on Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.5-1.png", + "caption": "Figure 1.5 The twin-disc machine.", + "texts": [ + " Two embodiments of what are known as EHL rigs are illustrated in Figs. 1.4 and 1.5. The most common today is the ball-on-disc machine shown in Fig. 1.4 where a steel ball is loaded against a disc which may be fabricated from steel or a transparent material such as glass or sapphire. When the disc is transparent, it is a simple matter to measure the film thickness of the liquid between the ball and disc by interferometry. However, there is a large variation in the accuracy of the various interferometry techniques. The twin-disc machine shown in Fig. 1.5 is better suited to traction measurement than film thickness; however, this configuration has been used for film thickness measurement as well by the electrical capacitance method. The shear force transferred across the film is known as traction or friction interchangeably. Traction in the rolling direction, longitudinal traction, may be measured as the force shown as the bold arrow in Fig. 1.4. The component of longitudinal traction due to shear of the liquid film results from a difference in velocities of the surfaces of the ball and disc in the rolling direction, say u1andu2", + " Traction is generally reported as a traction coefficient, the ratio of traction force to normal force or load, W. Equivalently, the traction coefficient is the ratio of average shear stress to average pressure, \u03c4=p. An additional kinematical condition is spin, the relative rotational velocity of the surfaces. Spin has been eliminated from the contact between the ball and disc in Fig. 1.4 by having the rotational axis of the ball intersect the axis of the disc at the contact plane. Other geometries may involve spin. Spin is generated by a nonzero tilt angle shown at the bottom of Fig. 1.5. Some measurements of central film thickness are shown in Fig. 1.6. The data points have been reported by INSA, Lyon [12 14] using their colorimetric interferometry technique that has shown to be very accurate and repeatable. The contact pressure here was low, pH 0:5 GPa, generated by the contact between a steel ball and glass disc. The film thickness of squalane, a low molecular weight (423 kg/kmol) branched alkane, shown as the squares, compares well to the theoretical Newtonian film thickness given by Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003348_icarm.2018.8610860-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003348_icarm.2018.8610860-Figure4-1.png", + "caption": "Fig. 4. The kinematics of the mobile platform", + "texts": [ + " The ZW axis is along vertical direction. The XW axis is along intersection line between vertical wall and horizontal plane. \u2022 Mobile base coordinate system (BCS) {B}: OBXBYBZB. According to the approximation above, the plane OBXBYB is always considered as being horizontal even though the mobile platform moves on an uneven ground. The origin OB locates on the bottom of the lifting column. The ZW axis is along vertical direction and the YB is parallel with the front surface of the column, as shown in Fig.4. \u2022 Lifting column coordinate system (LCS) {L}: OLXLYLZL.. The axis ZL is along the direction of movement of the lifting column. LCS may not coincide with the BCS which is caused by the incline of the ground. \u2022 Mounting plate coordinate system (PCS) {P}: OpXpYpZp.. It is fixed on the mounting plate which the manipulator arm is mounted on. PCS moves relative to the LCS along the axis ZL . \u2022 Manipulator coordinate system(MCS) {M}: OMXMYMZM.. It is the base coordinate of manipulator arm. MCS may not coincides with the PCS caused by the installation error between the manipulator and the mounting plate. \u2022 Tool coordinate system(TCS) {T}: OTXTYTZT.. It is fixed on the tool. The kinematics model of the mobile platform can be described by four parameters: xb, yb, zb, as shown in Fig. 4. Parameters [xb, yb, zb ] are the position of the base point in WCS. For the xb, yb, , they can be obtained by LMS laser range finder [14]. The parameter zb is the height of base point and it is affected by the uneven ground when the mobile platform moves. The vision system provides the value of zb., as shown in Fig.4. 978-1-5386-7066-8/18/$31.00 \u00a92018 IEEE 663 As shown in Fig.5, lifting column will tilt caused by the uneven ground or the manufacturing and assembly accuracy error of the mobile platform. The inclination sensor can obtain the pitch ( Pitch) and roll ( Roll) angles which are relative to the horizontal plane. PCS moves relative to the LCS along the column, H denotes the moving distance from OL. To be emphasized, the YM direction and ZM direction of MCS can hardly coincide with PCS completely because of the installation error between manipulator base and the mounting plate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002740_acoustics.2018.8502261-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002740_acoustics.2018.8502261-Figure1-1.png", + "caption": "Fig. 1. Test rig with asynchronous motor (1), metal bellow couplings (2,4), planetary gearbox (3), eddy current break (5), triaxial force sensors (6), accelerometers (7), gearbox mount (8) and piezoelectric inertial mass actuators (9).", + "texts": [ + " A modified adaptive feedforward control algorithm is proposed that is able to significantly reduce gear mesh vibration without worsening at other frequencies. Simulations are carried out to study the performance of the algorithm. Finally the performance of the algorithm is validated in experiments. To operate a gearbox for a geared turbofan engine isolated from the engine, a test rig with very large electric machines and a power circulation principle is required. To avoid this the AVC system is developed using a sub-scale test rig as depicted in Fig. 1. While the size of the gearbox is scaled down, the gear mesh frequency is kept constant. That is because the high frequency of the gear mesh is identified as one of the main challenges in the development of the AVC system. To keep gear mesh frequency constant while reducing the gearbox size and consequently the number of teeth, the small gearbox needs to run at higher input speeds 978-1-5386-7114-6/18/$31.00 \u00a92018 IEEE than the full-scale gearbox. In order to define requirements on the AVC system the gear mesh vibration of the selected planetary gearbox is measured and analyzed using the test rig", + " However passive gearbox mounts need to be modified to integrate the force sensors. In contrast to accelerometers they allow direct measurement of transmitted forces. Using this information, requirements for the anti-vibration actuators can be easily defined. In this work two piezoelectric inertial mass actuators are used for AVC. The gear mesh vibration can be suppressed on the test rig at frequencies up to 4.7 kHz. The test rig consists of an asynchronous motor that drives the sun gear and an eddy current break that is connected to the planet carrier as depicted in Fig. 1. Using the setup the gearbox can be operated at input speeds up to 10,000min\u22121 and output torques up to 40Nm. The gearbox mount is supported by two triaxial force sensors. In Fig. 1 the coordinate systems for both force sensors are sketched. Furthermore several mounting points for accelerometers are available at the gearbox mount. An operating point with an input speed of 10,000min\u22121 and an output torque of 30Nm is chosen for experiments as this is the full load test case with the highest achievable gear mesh frequency. The investigated planetary gearbox is commercially available and has a diameter of 60mm, 84 teeth on the ring gear, a weight of 0.9 kg and is spur-toothed. III", + " fmesh = Zfc \u2248 4666Hz (1) The frequencies of the neighboring sidebands are integer multiples of fc and can be calculated using Z \u00b1 1, 2, 3, ... in an analogous way. The nominal gear mesh frequency is almost invisible in the spectrum whilst the sidebands 83 and 85 are dominant with 60N and 18N respectively. However other sidebands from number 77 up to 89 exhibit also significant amplitudes. The goal of this work is to actively reduce the transmitted gear mesh forces in both z directions of the gearbox mount as they exhibit the highest vibration amplitudes. The piezoelectric inertial mass actuators labeled with (9) in Fig. 1 were developed in a previous work [12]. They consist of an inertial mass made of a small steel block that can be screwed onto a preloaded low-voltage piezoelectric stack actuator. The inertial actuator principle is selected because of its high flexibility. I n c ontrast t o a ctive m ounts, t he actuators do not have to be mounted in the load path, the original gearbox mount can remain unchanged. The properties of the actuator system are listed in Table I. The algorithm proposed in this work is depicted in Fig", + " Furthermore, the effect of the active vibration control onto transmission of gear mesh forces was studied in detail. The test scenario was again the described full load operating point of the gearbox. The phase-exact algorithm was implemented on a speedgoat performance real-time target machine featuring an Intel Core i7 3770K 3.5GHz quadcore CPU. The sampling frequency was set to 50 kHz. The algorithm was parameterized with step-size \u03b1 = 0.0008 and forgetting factor \u03b3 = 0.05. Nine sidebands from 81 to 89 were selected for cancellation. Two actuators were equipped as depicted in Fig. 1 to minimize the transmitted gear mesh forces in both z1 and z2 direction. For this purpose two SISO algorithms were implemented in parallel on the realtime target. A detailed study of the impact of the control onto the transmission of gear mesh forces was possible using the force orbits plots as introduced earlier in this paper. Fig. 5 depicts the force orbits for both sides of the gearbox mount in the uncontrolled and controlled case. Several effects are visible: The achieved reductions are summarized in Table II" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000671_pesc.2008.4592524-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000671_pesc.2008.4592524-Figure1-1.png", + "caption": "Fig. 1. la saturation effect and its sinusoidal spatial variation depending on the PM position.", + "texts": [ + " \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 \u2212+ \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 + \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 )cos( )cos( )cos( 3 4 3 2 \u03c0 \u03c0 \u03b8 \u03b8 \u03b8 m c b a ccbca bcbba acaba c b a s c b a \u03c8 i i i lMM MlM MMl dt d i i i r v v v (1) Such saliency is modelled as a sinusoidal spatial variation of the stator inductances [10] and mathematically these three stator inductances follow (2). ( )\u03b8\u22c5\u22c5\u2212= \u0394 Pllla 2cos0 (2a) ( )( )3 22cos0 \u03c0\u03b8 \u2212\u22c5\u22c5\u2212= \u0394 Plllb (2b) ( )( )3 42cos0 \u03c0\u03b8 \u2212\u22c5\u22c5\u2212= \u0394 Plllc (2c) Whenever the PM is close to any phase, the inductance in such phase will be saturated and therefore have a smaller value as Fig. 1 shows. This inductance value can be calculated by measuring the di/dt (by means of a non integrating Rogowski Coil [11]) caused by a known voltage pulse test as Fig. 2 illustrates. la= l0-l\u0394 cos (2\u00b790\u00ba)=l0+l\u0394 III. POSITION SIGNALS FOR MATRIX CONVERTERS Provided that this sensorless technique will be used at low and zero speed, the back-emf component of the machine model introduced in equation (1) might be negligible. Moreover, if voltage pulse test vectors are used, the machine can be further simplified to its high frequency model and the stator resistance voltage drop can be also neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000124_pime_auto_1964_179_013_02-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000124_pime_auto_1964_179_013_02-Figure2-1.png", + "caption": "Fig. 2. Velocities of the body c.g. Point 0,, and the c.g. are the body roll centre and the c.g. respectively. (A) Diagram illustrating that rotation about point 0, causes a linear velocity of the c.g. in a direction perpen-", + "texts": [ + "sagepub.comDownloaded from 100 J. R. ELLIS this is assumed to be at the height at which a line joining the front and rear spring attachment points to the body crosses a vertical line through the centre of the road wheel when the vehicle is viewed in side elevation. The point 0, may move vertically and horizontally. Rotation of the axle is assumed to take place about the centre of the axle, 02. The velocity of the c.g. of the body is made up of components of the vertical, lateral and rolling motiong, Fig. 2. vz = ( Z - H ~ sin ey+(i.-ad cos e l 2 or v2 = Zz+?2-2ZHd sin 0 - 2 e H d cos B+H2d2 The kinetic energy of the body is that due to the linear velocity plus that due to rotational velocity. Hence, Tbodv = $bfv2+*Id2 Substituting for v2 Tbody = f M ( Z 2 + p2-2ZHd sin 0-21;Hd cos 8+H2d2) Since the centre of rotation of the axle is assumed to coincide with the c.g. of the axle the kinetic energy may be expressed as the sum of the kinetic energies of translation of the c.g. in the vertical and horizontal directions plus the kinetic energy of rotation about the c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002089_iscaie.2018.8405443-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002089_iscaie.2018.8405443-Figure1-1.png", + "caption": "Fig. 1. CAD of cricket bowling machine.", + "texts": [ + " The corresponding x, y, and z components of the ball\u2019s translational and angular velocity were represented by (Vx ) , (Vy ) , (Vz ), \u03c9x, \u03c9y, and \u03c9z respectively. max= - 1 2 CD\u03c1A V .Vx- 1 2 CL\u03c1A V Vz\u03c9y-Vy\u03c9z \u03c9 (1) may= - 1 2 CD\u03c1A V .Vy- 1 2 CL\u03c1A V Vx\u03c9z-Vz\u03c9x \u03c9 (2) maz= - 1 2 CD\u03c1A V .Vz1 2 CL\u03c1A V Vy\u03c9x-Vx\u03c9y \u03c9 (3) V= (V1+V2) 2 (4) CL=3.19\u00d710-1[1- exp -2.48\u00d710-3\u03c9 ] (5) II. METHODOLOGY The performance analysis of the cricket bowling machine was divided into three test groups namely velocity, distance, and spin rate tests. The CAD of the fabricated machine used for the tests is illustrated in Fig. 1 and Fig. 2. 978-1-5386-3527-8/18/$31.00 \u00a92018 IEEE 47 A. Velocity test The velocity experiments were done on the machine to investigate the effects of two manipulated variables namely ball wetness and wheel pressure on the ball\u2019s translational velocity. The device used to measure the ball\u2019s initial velocity was a camera with high speed recording at 1000 frames per second placed at a distance on one side of the machine. On the opposite side, a black board with two vertical white lines was placed as the measuring scale" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003692_s1068799818040104-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003692_s1068799818040104-Figure3-1.png", + "caption": "Fig. 3. The bearing segment under radial load.", + "texts": [ + " This technique, obviously, neglects any ring deformations related to bending and some alternatives accounting properly this factor are to be presented further. Technique 2: Simplified piecewise deformation of a ring for equidistant distribution of unequal forces. For equidistantly distributed pointwise forces jQ let us consider that even if they are no longer equal, relation (9) still holds for any jth roller. Such an assumption, by fact, treats the ring as a set of segments of RUSSIAN AERONAUTICS Vol. 61 No. 4 2018 571 finite length R\u03b8 (Fig. 3) subject to a constant internal tensile/compressing force equilibrating the external radial loading. Distinctly from the first technique, this option provides a one-step solution stemming from system (4), (5). Further, we evaluate the limits of applicability of such a hypothesis. It is apparent that the introduced simplification, likewise the previously described technique, omits any deformations related to the ring bending. Technique 3: The ring deformed shape approximation with the fourier series expansion for an arbitrary loading case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure3-1.png", + "caption": "Fig. 3. First metamorphic limb with phase e1", + "texts": [ + " Phase e1 of metamorphic joint conform motion type of submanifold T2(w1) \u00b7U(p1,w1,u1) limb and phase of metamorphic joint coincide motion type of submanifold T2(u1) \u00b7 U(q1,u1,v1) limb combining the above constraint conditions. Therefore, transformation of motion type from submanifold T2(w1) \u00b7U(p1,w1,u1) to submanifold T2(u1) \u00b7U(q1,u1,v1) is correspond with transformation from phase e1 to phase e2 of metamorphic joint. 2) Type synthesis of first metamorphic serial limb: Submanifold T2(w1) \u00b7U(p1,w1,u1) limb has been generated by one of R(p11,w1) \u00b7R(p12,w1) \u00b7R(p13,w1) and R(p14,u1) with respect to local coordinate system o1 \u2212 u1v1w1. The metamorphic limb is synthesized by combining with the phase e1 in Fig. 3. Link a is fixed to base, and revolution joint of axis 2 is locked. Axis 1 rotate along with axis u1 direction. Axis 3 rotate along with w1 axis direction. Axis 4 and axis 5 are parallel to axis 3 with w1 direction. Then metamorphic limb takes on mobility four which can realize motion type submanifold T2(w1) \u00b7U(p1,w1,u1). Keeping rotational joint of axis 2 relaxed and rotational joint of axis 1 locked, rotate link c to axis 3 parallel with axis u1 direction in Fig. 4. Axis 2 rotate along with axis v1 direction, axis 4 and axis 5 are parallel to axis 3 with u1 direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001931_etechnxt.2018.8385373-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001931_etechnxt.2018.8385373-Figure6-1.png", + "caption": "Fig. 6. Flux characteristics plot presenting change in coenergy as a function of current and rotor position.", + "texts": [ + " 5, the calculated flux linkage characteristics are plotted with respect to current for different rotor positions. III. COMPUTATION OF TORQUE CHARACTERISTICS The torque characteristics can be achieved precisely by the use of a torque transducer. Due to the higher cost of the torque transducer, the use of torque transducer is avoided. In this paper, the estimation of torque characteristic is achieved using calculated coenergy. The mathematical expression for the calculation of coenergy is given later in this section. The change in coenergy equals to the area enclosed by the flux versus current. In Fig. 6, the shaded region 1 describes the coenergy change with respect to the rotor position. The net torque calculated is equal to the coenergy change with respect to the change in rotor position. The region 2 shows the coenergy change with reference from the linear flux model as shown by the lines OM and ON. The calculated coenergy and derived torque, are shown in Fig. 7 and Fig. 8. The torque profile is plotted from 0A to 55A in steps of 1A. The readings are recorded for each rotor position with a difference of 1\u02da between aligned and unaligned position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000640_icma.2009.5246170-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000640_icma.2009.5246170-Figure10-1.png", + "caption": "Fig. 10. Control motion for modified program motion at statically unstable regime (left), control motion for nominal program motion (right).", + "texts": [ + " The trajectories of the control movement are different for big and small errors integration. The proposed method of modifying the program movement is applied to strong instability regimes. For example it is investigated the movement of the machine at program motion by the low, showed at Fig.9. The walking of machine at this case is synchronous shift both the front legs and then the rear legs. This walking is analog of gallop walking. One walking is strong instability, because the lift legs from the same side is provide machine falling. At Fig. 10 (right) is presented the unstable control motion of the machine with falling. To stabilize the control motion is used the modified program motion. The modify of program motion at this regime have evident interpretation. That is appended longitudinal acceleration on locomotion direction, and so the result program motion is the movement with grew velocity - Fig. 9. The longitudinal acceleration has corrected negative vertical reactions at the fingers to positive. Note that longitudinal acceleration provided unlimited speed grows, and changes the parameters of robot walking. It method is produced the stability control locomotion - Fig 10. (left). The stable locomotion at that regime is depended from step length, for short step length doesn't provide the stable locomotion. The achieved results may be used at formulation the type of stability of the spatial complex mechanical system. It is necessary note, that the problem of stability mechanical system, described Eq. 1-3 can not be solved by the first Lyapunov method, because the system is strong nonlinearity. The asymptotic stability of the part of state variable of the Eq.I-3, defined the relative displacement and velocities at the drives is provide by proportional and differential feedback gains" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003119_s1063785018110196-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003119_s1063785018110196-Figure2-1.png", + "caption": "Fig. 2. (a) Schematic diagram of laser wire cladding: (1, 2) conical beams, (3) cladding material (wire), (4) focal spot on the wire, (5) focal spot on the substrate, and (6) substrate. (b) The dependence of radiation power that is sufficient to fuse the surface of the workpiece (upper curve) and surface of the tip of the wire (the lower curve) on the application rate.", + "texts": [ + " Second, the absorption of laser radiation increases up to the full absorption when the coherent part of the laser radiation is returned to the laser resonator by mirror 12. Laser wire cladding is a process that has many benefits over traditional clad processes. Usually, wire feeding is performed both from the side and coaxially when using a conical beam [4]. This approach uses focus spots of millimeter size and laser power at the level of a few kilowatts to heat both the substrate and the wire. The optimal heating of the wire and surface of the substrate is achieved when two conical beams (Fig. 2a) for separate heating of the substrate and wire are being used. \u03c0 \u03c1 = \u03c0 \u2212 \u22123 2(4/3) ( / ) ,ar C Dt dt r qQ E LG = \u2212 \u03c0 \u03c12/ /4 ,dr dt G r = 3 /4 .aT qtQ Cr NICAL PHYSICS LETTERS Vol. 44 No. 11 2018 The temperature distribution over the surface (z = 0) in the moving coordinate system XY for a moving source at a surface radiation absorption has the form [10] (7) where V is velocity, r is the radius of the focal spot, q is the power density, A is absorptivity, a is the coefficient of temperature conductivity, and k is thermal conductivity coefficient", + " The criterion for choosing of the solution is the ratio of exposure time te and time of heat conduction tc: (8) = \u03c0 + \u00d7 \u2212 + + 2 2 2 2 2 1/2 ( , ) [ /2 ( )] exp{[ ( ( ) )]/2 }, T x y Aqr k x y V x x y a = = 22 / , / .e ct r V t r a TECHNICAL PHYSICS LETTERS Vol. 44 No. 11 20 If tc > te and V > 2a/r, then the mode of fast motion is realized and the temperature distribution on the surface has the form (9) The calculations of the cladding process of 100-\u03bcm-diameter steel wire on a steel substrate were carried out. The diameter of the focusing spot on the substrate was 200 \u03bcm, and it was 80 \u03bcm on the wire (Fig. 2a). The calculations were carried out for the fast heating mode. The results are shown in Fig. 2b. The required capacity was much lower than at the standard modes of wire cladding when using a single laser beam. The use of coaxial laser cladding with several conical beams allows providing a lower rate of mutual dilution than in the case of other processes of cladding at high speed and with minimal heat input. An experimental system of laser cladding based on conical beams has been developed and manufactured (Fig. 3). The used optical scheme is shown in Fig. 3a. The possibility of heating a dense powder f low has been investigated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002870_rpj-11-2016-0178-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002870_rpj-11-2016-0178-Figure1-1.png", + "caption": "Figure 1 Sketches of the surface tessellation and tool path generation of assembled ABS specimens using two different assembling angles", + "texts": [ + " Tensile performance of the feed material was determined using Zwick Z010 machine equipped with a load cell capacity of 10 kN. Wires of a typical length of 108 mm and a diameter of 1.75 mm were tested. Testing was performed up to failure using a displacement rate of 5 mm/min. The design of 3D-printed parts taking the form of half-tensile specimens (length = 150 mm, width = 20 mm, thickness = 4 mm, gage area = 80 10 mm2) with an inclined edge was performed to study different balances between shearing and tension at the junction (Figure 1). The printing of 3D-printed specimens was performed using uPrint SE 3D printer from Stratasys (Figure 2). The built-in software (CatalystEX) was used to process all CAD (computer-aided design) files including the positioning of the samples in the base and the selection of appropriate printing parameters. The main process parameter was the assembling angle (a), which is the geometrical parameter that drives the modulation of interfacial behaviour. This angle was varied between 0\u00b0 and 90\u00b0 as shown in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002355_978-981-13-0629-7_5-Figure5.9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002355_978-981-13-0629-7_5-Figure5.9-1.png", + "caption": "Fig. 5.9 Flux density distributions under different flux density ratio c with optimal k when flux linkage of phase A is maximum", + "texts": [ + "4e, it can be noted that the FE obtained optimal values of split ratio are consistent with that from analytical results for both c = 0.5 and c = 0.8, as also can be seen from Table 5.3. Under c = 0.5 and optimal split ratio, the conventional has the highest average torque whilst the sine had the lowest. Under c = 0.8 and optimal split ratio, the torque of the sine+3rd is about 99.5% of the conventional. It can be concluded that the sine+3rd can be promising in torque density under higher flux density ratio for 12-slot/10-pole machines. Figure 5.9 shows flux density distributions for three machines with optimal k under c = 0.5 and c = 0.8 when flux linkage of phase A is maximum. It can be seen that the maximum flux density in stator lamination is around 1.5 T, which is below the knee point of lamination steel B-H curve, 1.6 T. The saturation level of the machines is low. 5.5 Experimental Verification The 12-slot/10-pole SPM machine with sine+3rd-shaped rotor has been prototyped and tested for validation. The stator and the sine+3rd-shaped rotor are pictured in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002483_aim.2018.8452437-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002483_aim.2018.8452437-Figure11-1.png", + "caption": "Fig. 11: Overview of experimental setting for wheel bulldozing test.", + "texts": [ + " However, this effect relatively small. Therefore, the differences of the value between experiment and simulation result are relatively small. These results confirm that the Hegedus\u2019s model is able to represent the bulldozing force of the plate on Silica sand No. 5. In order to measure bulldozing force of backward wheel, a wheel bulldozing experiments werer conducted in the same way as the experiment of the metal plate. 1) Experimental environment and conditions: The overview of the experimental system and environment are shown in Fig. 11. The bulldozing area and soil are same as the plate bulldozing experiment. The bulldozing force is obtained by force sensor, which set up upper of the wheel. The rope, which connects to the pole, pulls the wheel. In order to confirm the applicability of Hegedus\u2019s model to a wheel, the wheel size is set three sizes. The sizes are 150 [mm], 170 [mm] and 200 [mm] diameter and 40 978-1-5386-1854-7/18/$31.00 \u00a92018 IEEE 933 [mm] width. The wheel of 170 mm diameter is the same wheel size of the rover, which equips with advanced scheme (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000145_12.803180-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000145_12.803180-Figure2-1.png", + "caption": "Fig. 2 Bearings: a - rolling, b \u2013 tilting-pad, c \u2013 sleeve liquid-friction", + "texts": [ + " Different technological machines are used in different ranges of industry and are worked with different characteristics being different conditions of work. Parameters of work of technological machines are depended from different factors: from frequency of rotor rotation, from temperature of work, from vibrations and so on. Subject to type of using bearings are depended vibrations of rotor, accuracy of rotor rotation and so on. Researches of accuracy of rotor rotation are done using technological machines (Fig. 1) with segmental and sleeve liquid-friction bearings of sliding friction and bearings of roll (Fig. 2). Bearings are one of the primary elements of technological machines, they are kept rotor and are pervaded in different systems of rotor. Bearings of roll are used in little and mean power machines: in electric motors, in generations, in centrifugal pumps, in ventilators and so on. Bearings of sliding friction are used: in turbines of vapour of large power, in compressors, in pumps, in generators of internal-combustion, in generators and similarly. Eighth International Conference on Vibration Measurements by Laser Techniques: Advances and Applications, edited by Enrico Primo Tomasini, Proc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002082_1.4040813-Figure19-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002082_1.4040813-Figure19-1.png", + "caption": "Fig. 19 Pressure profiles versus oil supply flow rate, 10,000 rpm, 365 kPa specific load", + "texts": [ + "org/about-asme/terms-of-use less pronounced in the flooded model predictions than the starved model predictions where it is quite significant. These trends suggest that the models are most likely overpredicting bearing damping. Unlike the nominal load condition, oil supply flow rate has a very marginal effect on the predicted damping ratios of the starved model at the higher, 365 kPa load condition as shown in Fig. 10. The starved model still, however, predicts greater damping ratios and system stability than the flooded model under these conditions. Figure 19 illustrates the predicted pressure profiles of the starved bearing model for varying oil supply flow rates under the increase bearing load condition at 10,000 rpm. For the range of flow rates tested in this study, oil supply flow rate has no noticeable effect on pad pressures. As expected from these results, predicted stiffness and damping coefficients vary little with oil supply flow rate and are not presented here. The relatively large decreases in bearing stiffness and damping with decreasing flow rate predicted by the starved model in the lower load case are nonexistent under the higher load condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002860_s11668-018-0561-y-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002860_s11668-018-0561-y-Figure1-1.png", + "caption": "Fig. 1 Heat treatment drawing of drive axle, showing the profile of induction-hardened case", + "texts": [ + "\u2019 In an attempt to verify this hypothesis, a simulated test was conducted which confirmed that delayed cracking could occur if inductionhardened axles were kept un-tempered overnight at low ambient temperatures. Another interesting and rare feature of the present fracture was that the fatigued area was \u2018curved,\u2019 i.e., the fatigue crack, as it progressed, had gradually acquired an orientation that was normal to the maximum operative tensile stress. Keywords Induction hardening Quenching stresses Delayed cracking Dye penetrant inspection Intergranular cracking Wheel axle (also called \u2018Stub Axle\u2019), as shown in Fig. 1, supports and drives the rear wheels of the tractor. In Massey Ferguson (Pakistan) tractors, this component is made from 080-A47 steel by forging, and that prior to machining, the forgings are normalized to a Brinell hardness of 217 (max) and a grain size equivalent to ASTM size 6\u20138. After machining, the axles are induction-hardened to achieve a case profile in compliance with the drawing shown in Fig. 1. The induction hardening treatment is carried out on a 200 kw/10 kHz machine using a programmed power and dwell control to achieve the desired case profile. The present failure incidences pertain to a production lot of about 4000 axles manufactured during November 2016\u2013 January 2017. As many as 49 axles broke during service in almost identical pattern and, as per the warranty policy of the company, were replaced along with the cost of consequential damage. The broken components were returned to the tractor manufacturing facility at Lahore (Pakistan) for a detailed investigation of the failure", + " Initial crack, which (as discussed later) was essentially within the induction-hardened case. 2. Progressive fatigued region, showing typical \u2018beach marks.\u2019 3. Final rupture which appears to be essentially shear. The depth of the \u2018origin\u2019 of the crack in almost all the broken axles was observed to be about 3\u20134 mm. Some of the broken axles were longitudinally sectioned and macroetched to examine the profile of induction-hardened case. Figure 4 shows the profile of hardened case, which was found to be in compliance with the specification as shown in Fig. 1. Figure 4 also shows that the original crack had formed in the \u2018neck\u2019 region of the axle and that the depth of the initial crack (which is clearly visible in Fig. 3) was essentially the same as the depth of the induction-hardened case. It was also noticeable that this initial crack was normal to the surface at the location where it was formed, a feature which is typical of quench cracking [1]. The \u2018curved\u2019 path of the fatigue crack may also be noted. Microstructure shown in Fig. 5, which was taken from a location indicated by arrowhead, shows that the crack was normal to the surface and was intergranular, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000983_icelmach.2008.4800111-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000983_icelmach.2008.4800111-Figure3-1.png", + "caption": "Fig. 3. Phasor diagram of PMSG connected to R-L type load.", + "texts": [ + " This paper describes the implementation of loading method for the calculation of load characteristics of stand-alone PMSGs operating at constant rotational speed and supplying the R or R-L type loads. The field-circuit loading method was at first developed for the calculation of electromechanical characteristics of permanent magnet synchronous motors (PMSM) [1] \u00f7 [3]. In this method, the electromechanical characteristics of PMSM at steady-state operation are essentially predicted basing on the classical equations of a salient pole synchronous machine derived from a model of this machine in the d-q reference frame. The lumped parameters of the d-q phasor diagram (see Fig. 3) which have the most important impact on the performance of synchronous machine: d-axis magnetizing reactance Xmd, q-axis magnetizing reactance Xmq and opencircuit voltage E0 are in the loading method calculated separately for each particular point of machine operation (load condition) with the use of magnetostatic, two-dimensional finite element analysis (FEA) of magnetic field distribution [1]. The variations of Xmd, Xmq and E0 with change of load current and magnetic saturation level are taken into account", + " The main idea of loading method adopted for the calculation of load characteristics of stand-alone PMSG is exactly the same as in the case of loading method used for the prediction of electromechanical characteristics of PMSM. It is based on the possibility of calculation (after previous FEA of field distribution inside the machine) of an internal voltage Ei and internal power angle \u03b4i values, corresponding to the actual field distribution at each particular machine operating point, assuming that a modulus I1 and a phase angle \u03b3 of an armature current phasor I1 are known \u2013 see Fig. 3 and Fig. 5. A model of PMSG for FEA is prepared so that the A-phase axis of stator winding is exactly aligned with the d-axis of rotor \u2013 see Fig. 4. Additionally, we assume that the FEA model corresponds always to an instant of time t = 0, and that the instantaneous values of phase currents, which are the input variables for FEA, are described by equations: In such a case, appearing in (2) the phase angle \u03b2 of the stator current phasor I1, calculated with respect to the rotor d-axis: \u03b2 = \u03c0/2 \u2013 \u03b3, corresponds simultaneously to the space angle between a stator magnetomotive force (MMF) vector Fs and a rotor MMF vector Ff (Fig", + " Next, the modulus Ei of the internal voltage phasor Ei and the internal power angle \u03b4i \u2013 the phase angle of the resultant air-gap flux vector (see Fig. 5) are calculated using (4) and (5): sui kkzfE \u22c5\u22c5\u22c5\u22c5\u22c5= 11144.4 \u03a6 , (4) ( )11 abarctgi =\u03b4 . (5) In equation (4) f1 is the frequency of stator current, z1 is the number of winding turns in series per phase, ku1 and ks are the winding distribution factor and the slot skew factor respectively. Immediately after calculation of Ei and \u03b4i we can calculate the below listed quantities using equations (6) \u00f7 (15) delivered mainly from the phasor diagram shown in Fig. 3: \u2022 the modulus V1 of the phase voltage phasor V1 (and then the modulus VLL of line to line voltage at machine terminals): 111 RXjV \u22c5\u2212\u22c5\u2212= 11 IIEi , (6) \u2022 phase angle arg_V1 of the voltage phasor V1: ( )111 arg RXjarg_V \u22c5\u2212\u22c5\u2212= 11 IIEi , (7) \u2022 power angle \u03b4: 12/ arg_V\u2212= \u03c0\u03b4 , (8) \u2022 electromagnetic (internal) power Pei: ( )iiiei IIEP \u03b4\u03b2\u03b4\u03b2 sincoscossin3 11 \u22c5\u22c5\u2212\u22c5\u22c5\u22c5\u22c5= , (9) \u2022 electromagnetic torque Te: ( )nPT eie \u22c5\u22c5= \u03c030 , (10) \u2022 mechanical power at the generator shaft P1: mei PPP \u2206+=1 , (11) \u2022 active electrical power delivered by the generator P2: \u03d5cos3 12 \u22c5\u22c5\u22c5= IVP LL , (12) \u2022 reactive electrical power delivered by the generator Q: \u03d5sin3 1 \u22c5\u22c5\u22c5= IVQ LL , (13) \u2022 apparent electrical power delivered by the generator S: 13 IVS LL \u22c5\u22c5= , (14) \u2022 efficiency \u03b7: 12 PP=\u03b7 ", + " From this additional magnetic field computation internal voltage Ei\u2019 and internal power angle \u03b4i\u2019 are obtained. After that we can calculate the lumped parameters of equivalent circuit Xmd and E0 appropriate for a specific load condition: \u03b2\u03b2 \u03b4\u03b4 coscos coscos ' 11 '' \u22c5\u2212\u22c5 \u22c5\u2212\u22c5 = II EE X iiii md , (17) mdii XIEE \u22c5\u22c5\u2212\u22c5= \u03b2\u03b4 coscos 10 . (18) In the developed calculation algorithm, based on the fieldcircuit loading method, to solve the set of equations derived from the phasor diagram of stand-alone PMSG connected to R or R-L type load (Fig. 3) it is assumed that the input (known) variables are: the phase angle \u03b3 (so also the phase angle \u03b2) of the current phasor I1 and the load power factor cos\u03d5. With given values of \u03b3 and cos\u03d5, all of the equations describing the stand-alone operation of PMSG with R or R-L type load are satisfied simultaneously by just one value of the armature current modulus I1. The correct value of modulus I1 is not known in advance and it must be found using iterative estimation methods. In the developed algorithm these iterative calculations are realized with the secant method. The iterative solution \u2013 correct value of modulus I1 \u2013 is found if the value of phase angle of the voltage phasor V1 calculated using (7) is equal to value calculated as the sum of angles \u03b2 and \u03d5 (see Fig. 3 and Fig. 5). For each successive step of iterative calculations for finding appropriate value of I1 one nonlinear FEA of magnetic field distribution inside the generator is carried out. Other principal assumptions taken into account in the developed algorithm (and software) for the calculation of load characteristics of stand-alone PMSG with R or R-L type load can be summarized as follows: \u2022 to derive the load characteristics for a range of output powers (e.g. VLL = f(P2)), calculations are realized for a given range of phase angles \u03b3 of the armature current phasor I1; \u2022 the load power factor cos\u03d5 is given in advance and it is constant in the entire range of considered generator loads; \u2022 for each particular value of armature current phase angle \u03b3 there is only one value of armature current modulus I1 that satisfies simultaneously all the generator equations derived from the phasor diagram shown in Fig. 3 (the increase of phase angle \u03b3 causes the increase of armature current modulus I1). To verify the correctness of developed algorithm and software for computation of stand-alone PMSG load characteristics, these characteristics were calculated for permanent magnet synchronous machine type PMSg132S4. It was assumed that this machine was operating as stand-alone generator connected to R-L type load with power factor cos\u03d5 = 0.99. Machine type PMSg132S4 is of surface mounted construction, the magnetizing reactances in d and q rotor axes, Xmd and Xmq, have very similar values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003016_j.compstruct.2018.11.095-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003016_j.compstruct.2018.11.095-Figure7-1.png", + "caption": "Fig. 7. (a) Illustration of the use case surface with a representation of the Gaussian curvature (blue colorization) (3) for maxium curvature. (b) Graph with algorithm results (mean deviation, starting orientation and maximum overlap) for the first algorithm stage (iteration step 1\u2013100) and for the second stage (iteration steps 101\u2013200).", + "texts": [ + " For each iteration step the fitness criterion dev mean step, , is exported. In addition the greatest determined overlap is exported for discussion of the results as the maximum overlap distance lapmax . For application of multiple tapes in an automated process, the middle trajectories (of the best solution) are provided as a polygon. The layup head is able to laterally tailor the tapes (see Section 4). Thus, the edge trajectories for each tape are calculated to enable a constant minimum overlap lapmin. The surface which has to be covered by multiple tapes (see Fig. 7(a)) represents the double-curved region with geometric transition in the cylindrical fuselage section. The curved surface is trimmed with two planes, which have an offset of 3600mm. This results in a border edge with the dimensions: W1 = 2640 mm, W2 = 910 mm, L1 = 3630 mm and L2 = 3930 mm. In the old multiple tape arrangement, this surface will be covered with 6 tapes of =W 630 mmT . The enlarged surface (2) enables calculation of start and end of the tapes. The coloring represents the Gaussian curvature with its maximum =G 1", + " In order to find an application-driven solution, we set the fixed parameters as follows:WT = 800 mm, LT = 2800 mm, lapmin = 30 mm and lapmax = 600 mm. We assume a local minimum of the fitness criterion dev mean step, , approximately for = \u00b00dev start T, , 1 . Therefore, we focus the search on = \u00b1 \u00b017.5dev start T, , 1 , so that = \u00b00dev start T, , 1 is given for step 50. 100 iteration steps are performed for each phase of the twophase local search. Hence, the increments are set to = \u00b00.35phase1 and = \u00b00.0075phase2 . Fig. 7(b) displays the fitness criterion dev mean step, , (black squares) and the maximum overlap distance lapmax (blue triangles) for first tape orientation dev start T, , 1 (red circles) and every iteration step. In the first 18 iteration steps, at least one tape left the extended surface. Therefore, no solution was calculated. In phase 1 at iteration step 72, the minimal fitness criterion = \u00b09.27dev mean, ,72 is calculated for = \u00b07.7dev start T, , 1,72 . The maximum overlap distance results in =lap 141 mmmax,72 ", + "24dev mean, ,178 and =lap 139 mmmax,178 for = \u00b07.935dev start T, , 1,178 . Each calculated solution covered the surface with 5 aligned tapes. In addition to the deviation output of the algorithm, the geometrical information is generated for each pattern. The multiple trajectory geometric plot of two solutions with tapes is shown in Fig. 8. Here, an exemplary not optimal solution (iteration 43) on the left and the best solution (iteration 178) on the right are shown. The green dotted line represent the edge of the surface (1) see Fig. 7(a). The black arrow illustrates the starting orientation of the first tape, and the red arrows the maximum deviation dev max, for every single tape on the surface. The graphs (Fig. 7) and plots (Fig. 8) give information on the algorithm for multiple tape designs. The results in the graph (Fig. 7(b)) show that the mean deviation dev mean, is not directly dependent on the starting orientation dev start T, , 1 of the first tape. The mean deviation is partially erratic. The non-deterministic local search is thus advantageous for finding suitable solutions. In addition, it is clear that a minimum of the desired mean deviation (iteration step 72) is not achieved when the first tape is perfectly oriented at = \u00b00dev start T, , 1 (iteration step 50). However, since we reasonably limited the optimisation parameters it is not ensured that it is the global minimum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001813_978-981-10-8306-8_13-Figure13.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001813_978-981-10-8306-8_13-Figure13.5-1.png", + "caption": "Fig. 13.5 The Tresca p diagrams of a the yield surface and b the plastic strain increment surface", + "texts": [ + "22) is replacable with the following Hosford effective stress: r \u00bc a SI SIIj j \u00fe SII SIIIj j \u00fe SIII SIj j\u00f0 \u00def g \u00f013:35\u00de Plot the p diagram for the Tresca yield surface by applying Eq. (13.35) and while considering the stress distribution shown in Fig. 12.6. Also, derive that a \u00bc 1 2 or a \u00bc 1 4, respectively, when the reference state is a simple tension (and balanced biaxial) state or a pure shear state. The Tresca p diagrams of the yield surface and the plastic strain increment surface are plotted based on the dual normality rules in Fig. 13.5. Note that there are multiple normal directions at the corners, which are obtained by considering a smooth curve that converges at each corner as its limit as shown in Fig. 13.6. Therefore, plastic strain increments are not unique for the simple tension or balanced biaxial stress states. Also, for the pure shear stress state, the plane strain deformation is obtained, for which one of three plastic strain increment components vanishes. However, there are multiple stress states for plane strain deformation from the simple tension to the balanced biaxial", + " For c \u00bc 0:5, the deformation becomes equivalent with that of the balanced biaxial state for the smooth incompressible and isotropic yield function. For c \u00bc 0:0 and c \u00bc 1:0, the deformation is the plane strain deformation, which is also obtainable with the pure shear stress state as shown in Fig. 13.7a. The figure confirms that depBBIII \u00bc depPSI \u00bc dB \u00bc 1 2 dK \u00f013:39\u00de Which complies with Y = B = 2K, as confirmed by the plastic work equivalence principle. The p diagram of the plastic strain increment surface shown in Fig. 13.5b implies that the conjugate Tresca effective plastic strain increment is d e \u00bc b deIj j \u00fe deIIj j \u00fe deIIIj j\u00f0 \u00def g \u00f013:40\u00de as the p diagram of Eq. (12.35) with M = 1.0 shown in Fig. 12.20 suggests. Plot the p diagram of Eq. (13.40), which should be the same as that shown in Fig. 12.20. Also, derive that b \u00bc 1 2 for the simple tension and balanced biaxial reference states and b \u00bc 1:0 for the pure shear reference, considering Eqs. (13.36) and (13.38). Plastic strain increments of Drucker and modified Drucker yield functions For the Drucker and modified Drucker yield functions, their conjugate plastic strain increment functions are not available for their analytical expressions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002514_msf.931.3-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002514_msf.931.3-Figure3-1.png", + "caption": "Fig. 3. The diagram of wind pressure and its approximation by means of the Fourier series", + "texts": [ + " For each n, a solution of a linear algebraic equations system is performed, having the form: { } { }.n n nK F\u03b4 = (13) Further, the displacements and internal forces are summed over each harmonic. Formulation of the Problem We consider a cylindrical tank, subject to wind pressure (Figure 2). Experimental studies (blowing of cylinders) show that for smooth cylindrical shells, the wind load acts perpendicular to the surface of the cylinder and changes in a circular direction according to the law depicted in Figure 3a [7]. A similar distribution of wind pressure is adopted in the normative literature [8]. Depending on the degree of blown cylinder surface roughness, the pattern of wind pressure distribution somewhat changes. For a rougher surface, the transition point from pressure to suction moves slightly towards the wind. The wind load components in the meridional and circumferential direction are assumed to be zero, and the normal component of the wind pressure can be represented by the Fourier series [7]: 0 cos . k n n W W n\u03b8 = =\u2211 (14) As experiments show, when function of wind pressure on a cylindrical surface is approximated, one can confine oneself to a series of three terms (Figure 3.3, b), ie, to set the function W(\u03b8) in the form: ( ) ( ) 2 0 0 cos 0.7 0.5cos 1.2cos 2 ,n n W W n p\u03b8 \u03b8 \u03b8 \u03b8 = = = \u2212 + +\u2211 (15) where p0 \u2013 maximum value of wind pressure. When calculating the corrugated shell, we assume that the wind pressure distribution for it is the same as for the smooth one. Results and Discussion A calculation was made for a hingedly fixed at the base shell, the radius of which varied according to the sinusoidal law: 0 2sin ,zr R A \u03c0 \u03bb = + (16) where R0 is the average shell radius, \u03bb is the wavelength, and A is the wave amplitude" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000428_ssp.144.175-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000428_ssp.144.175-Figure4-1.png", + "caption": "Fig. 4. Components of a Feed Drive of a CNC Machine Slide", + "texts": [], + "surrounding_texts": [ + "The machine structure is the load carrying and supporting member of the machine tool. All the motors drive mechanisms and other functional assemblies of machine tools are aligned to each other and rigidly fixed to the machine structure. The machine structure is subjected to static and dynamic forces and it is, therefore, essential that the structure does not deform or vibrate beyond the permissible limits under the action of these forces. All components of the machine must remain in correct relative positions to maintain the geometric accuracy, regardless of the magnitude and direction of these forces. The machine structure configuration is also influenced by the considerations of manufacture, assembly and operation. The following are some of the important constituent parts, and aspects of CNC machines to be considered in their designing [3]. a) Machine structure, b) Guide ways, c) Feed drives, d) Spindle and spindle bearings, e) Measuring systems, f) Controls, software and operator interface, g) Gauging and tool monitoring. Calculation of Drive Requirements on Feed Motor Shaft For the selection of an optimum servomotor and drive, it is necessary to find the drive requirements on the motor shaft. These are as below: a) Load Torque, b) Load Inertia, c) Maximum Speed. Generally, the following procedure is followed to find the load torque, load inertia and maximum speed requirements on the motor shaft and then it is selected proper motors in according to total inertia and maximum speed of table at the cutting process [3, 4, 5]. Main parts of CNC milling machine has been shown figure 2 and Fig. 3. It was used solid works to design the CNC machine as a detailed and It has been observed all movement the axes on the simulation process of the program before the producing the CNC machine. Some values can be taken from product catalogues, some of them can be taken from machine design or machine cutting books and some can be calculated by using concerning equations and constant values of machine tools and machining materials Torque for cutting at maximum speed of table T C = ( Ft.p B )/(2\u03c0 R\u03b7 ) =( 1100\u22c50.005)/( 2\u03c0 \u22c51 0.9 )= 0.97 Nm Torque for sliding mass T L=(W + f ) P B /(2\u03c0 R\u03b7 )=(200+ 80)\u22c50.005/(2\u03c0 \u22c51\u22c50.9)=0.25 Nm Total torque for sliding mass and cutting process T T = T L + T C = 1.22 Nm Inertia of coupling J C = (WcDc 2 )/8 = (0.2\u22c50.050 2 )/8 = 6.25\u00d710 -5 kgm 2 Inertia of ball screw Jb= (\u03c0 \u03c1 L B D B 4 )/3=(\u03c0 \u22c57.85\u22c510 3 \u22c50.5\u22c50.02 4 )/32=6.16\u00d710 5\u2212 kgm 2 Inertia of sliding mass, J s = W(P B /2\u03c0 R) 2 = 200(0.005/2\u03c0 \u22c51) 2 = 0.12\u00d710 -5 kgm 2 Total load inertia of moving parts J L = J 1.L + J B + J C = 12.53\u22c510 -5 kgm 2 Acceleration power of motors P a = (2\u03c0 N M /60) 2 .(J L /t a ) = 124 W Required Motor Power for cutting P 0 = ( 2\u03c0 N M T T )/60 = 383 W Required motor power for cutting is 383 Watt. Selected motor power is 400 Watt which has been shown on table 1. Therefore there is not any problem in cutting process of machine parts. X-Y and Z axes motors have been directly mounted to ball screws. It has been used trigger belt on the mounting spindle motor to spindle. Spindle speed has been decreased 3 times by using 3:1 rated trigger belt. Torque of motor should be satisfactory to cut the machine parts on the X-Y-Z axes." + ] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure4-1.png", + "caption": "Fig. 4 Trumpet robot lung", + "texts": [ + " As can be seen in Fig. 2, trumpets have three valves. The combinations of valves change the length of tube as in Fig. 3. Thus, the player needs to control the pitch of a sound both by use of three valves and by adjusting the frequency of lip vibration, and adjust the magnitude of sound by changing the volume of exhaled air. Imitating this mechanism, we developed both artificial lips and lung that could produce clear sound of the trumpet. The lung is designed compact enough to be installed in a robot as shown in Fig. 4. The hand is designed to have the ability to open and close piston valves quickly enough (Fig. 5). Figure 6 shows the software system for the trumpet performance. Both the sound pressure \u2013 to be precise, the vibration of the mute attached to the trumpet \u2013 and the valve operation have been recorded during human musical performance. The system converts the recorded sound pressure data into the pattern of the artificial lips vibration. Finger motion pattern is generated from the recorded valve data or reference MIDI data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003692_s1068799818040104-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003692_s1068799818040104-Figure1-1.png", + "caption": "Fig. 1. Bearing coordinate system and loading.", + "texts": [ + " In the present work, the following techniques are developed to incorporate structural flexibility of the rings, the shaft and the housing: 1) tensile (compressive) deformations from equidistantly distributed equal radial forces; 2) simplified piecewise deformation of the ring for equidistant distribution of unequal forces; 3) the ring deformed shape approximation with the Fourier series expansion for an arbitrary loading case; 4) structural FE assembly with specially designed nonlinear elements modeling rollers for an arbitrary loading case. The use cases of the presented formulation are discussed and their pros and cons are highlighted. A numerical benchmark test is performed to demonstrate the performance of the models. Consider a bearing as an element with two nodes iO and eO attached to centers of the inner and outer rings, respectively (Fig.1). The node eO is fixed and has the global cartesian system x , y , z defined. Subject to external forces = T x yF F\u23a1 \u23a4\u23a3 \u23a6F (1) the inner ring exhibits small deflections = T x yu u\u23a1 \u23a4\u23a3 \u23a6y (2) with respect to the node eO . The model is not capable to undergo axial loading, so the displacement component zu is omitted. Possible misalignment of rings is also neglected. RUSSIAN AERONAUTICS Vol. 61 No. 4 2018 569 Radial deflection of the inner ring in the coordinate system attached to the j th rolling element with the angular position = ( 1)j j\u03c8 \u2212 \u03b8 , where 2 = N \u03c0 \u03b8 and N is the number of rollers, can be determined as = ,rju jS y (3) where = cos sinj j j \u23a1 \u23a4\u03c8 \u03c8\u23a3 \u23a6S " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002882_1.5080567-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002882_1.5080567-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the launching apparatus, the projectile trajectory, and some of the relevant variables.", + "texts": [ + " Many introductory texts1 show that for a launch speed v, a launch angle \u03b8 relative to the horizontal, and a uniform gravitational field of magnitude g, the horizontal range R of a projectile, ignoring all forms of friction or drag, is given by (1) This range formula, as well as the one derived below, takes R to be the horizontal component of the displacement that brings the projectile back to its original launch height. A recent derivation of this result using differential calculus was presented by Allen.2 Following the approach of Schnick,3 we can determine v for a given launch mechanism from conservation of energy. For a projectile of mass m launched from a massless spring with spring constant k that has been compressed by an amount x from its equilibrium length (see Fig. 1), we can write (2) It is assumed here that an element of the launcher brings the spring to an abrupt stop when it reaches its equilibrium length, so that the spring does not apply a force to the projectile beyond that point. THE PHYSICS TEACHER \u25c6 Vol. 56, December 2018 585 Fig. 3. Comparing the two plots, we conclude that the projectile weight has the effect, via the reduction in the spring acceleration and the launch speed, of decreasing rmax by a small amount. Subtracting Eq. (6) from Eq. (9), we can calculate that the difference approaches \u221a2 as f grows large and \u03b8max approaches 45\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002175_978-3-319-99262-4_2-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002175_978-3-319-99262-4_2-Figure1-1.png", + "caption": "Fig. 1. Thrust bearing (a) variables and (b) oil film thickness profile", + "texts": [ + " The assumptions admitted on this work are laminar flow in thermo-hydrodynamic steady state operation, with no cavitation and no slip at fluidsolid interfaces; the inertia and body forces are negligible compared to viscous and pressure terms in the momentum equations; the lubricant is a Newtonian fluid, with constant density, thermal conductivity and specific heat and its viscosity is a function only of temperature; the thermal and elastic distortions are neglected (the surface pads and runner are rigid); and the velocity gradients across the film are much more important than all other velocity gradients, because of the film thin thickness. The geometry of the bearing considered in the simulations is schematically shown in Fig. 1. The oil film shape may be specified a priori, since no thermal and elastic deformations are considered in the model. The film thickness for the pad is shown in Eq. 1. h r; h\u00f0 \u00de \u00bc h0 \u00fe sh 1 h hramp ; h hramp h0; h[ hramp ( \u00f01\u00de wherein h0 is the minimum film thickness, sh is the shoulder height and hramp is the angular length of the converging gap, schematically shown. To study the thrust bearings, it is more convenient to write the governing equations in cylindrical coordinates. It is common practice to neglect the pressure variation along the film thickness, so the pressure p is only function of the radial and circumferential coordinates, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002162_978-3-319-96181-1_3-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002162_978-3-319-96181-1_3-Figure1-1.png", + "caption": "Fig. 1. Components of a rolling element bearing.", + "texts": [ + " Starting from the analysis of signals collected by vibration sensors of the bearing test rig, with the calculation of time indicators (kurtosis, Rms, or crest factor) and frequency indicators. Then, configure them to build the database which will be used for learning and testing the ANN, which will allow us to find the best network configuration (inputs, outputs and parameters), and subsequently to automate the decision on the possibility of the fault bearings. The main components of rolling bearings are the inner ring; the outer ring, the rolling elements, and the cage (see Fig. 1). Typically, the inner ring of the bearing is mounted on a rotating shaft, and the outer ring is mounted in the stationary housing. The rolling elements may be balls or rollers. The balls in a ball bearing transfer the load over a very small surface (ideally, point contact) on the raceways (Randall and Antoni 2011). Local or wear defects causes periodic impulses in vibration signals. Amplitude and periodic of these impulses are determined by shaft rotational speed, fault location, and learning dimensions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002493_s11277-018-5949-1-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002493_s11277-018-5949-1-Figure5-1.png", + "caption": "Fig. 5 Spatial node relations for nodes and based on distance and angle", + "texts": [ + " The representation of the regional coordinate system creates the spatial landmark relations as follows. 4. In this case, the robot makes spatial landmark relations about the recognized landmarks around it. If no other landmarks are observed, then the semantic mapping is complete with the current node. Figure\u00a03 shows the results, which can be expressed as 5. As indicated by (d) in Fig.\u00a04, the robot approaches to the new landmark to estimate distance and angle. If the landmark did not find on a previous memory, a new node is established, as showed by the regional coordinate system for (e) in Fig.\u00a05. 6. The robot creates spatial landmark relations of the landmarks at position (e) in Fig.\u00a04. If no more landmarks appear, the robot has completed its semantic-metric map with the current node. { \u0398 , S } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v )}} 1\u0393 = { \u0398 , S , } a,b\u22081O{ \u0398 , S , } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v ) , ( 1 , 1 )}} { \u0398 , S , } = { ( 1 , 1v ) , { ( 1 , 1v ) , ( 1 , 1v ) , ( 1 , 1 )}} 2\u03a9 = \u27e82v, \u27e9 \ufffd \u0398 , S \ufffd = \ufffd \ufffd 2 , 2v \ufffd , \ufffd \ufffd 2 , 2v \ufffd , \ufffd 2 , 2v \ufffd\ufffd\ufffd 1 3 7. Spatial node relations are made from the previous node to the current one as follows: 1,2s v = (||| 1X \u2212 2X ||| ) , 1,2s v = ( \u2220 ( 1X \u2212 2X )) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002790_detc2018-86262-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002790_detc2018-86262-Figure4-1.png", + "caption": "FIGURE 4. CROSS PIVOT FLEXURE IN DEFLECTED POSITION, DESIRED DOF IN RED, ACTUAL IN GREEN", + "texts": [ + " For both wrenches the linear combination vectors \u03bb have to be obtained. At the neutral position there was no parasitical translation. This can be seen in Figure 3 where the axis roll is zero at the neutral position. However, there was a parasitic rotation, which is indication by the axis shift. The leaf spring flexure in the previous example loses its translational freedom and gains a rotational degree of freedom if the leaf springs are crossed. All parameters are kept the same in this example, however the leaf spring flexures are angled 45 degrees (Figure 4). In the neutral position the flexure mechanism has the following induced twists, eigentwists, and eigencompliance. 6 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 12/09/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use T\u0302 = [ T\u0302\u03c1 T\u0302f ] = \u22120.0265 0.0075 0.0075 0 0 1 0 \u22120.0265 0.0265 1 0 0 0.0265 \u22120.0075 \u22120.0075 0 1 0 0 0 1 0 0 0 1 0 0 0 0 0 0 0 0 0 0 0 (29) c\u03b5 = diag(0.48 0.001 0.001 2.5\u00b710\u22127 2.1\u00b710\u22128 2.1\u00b710\u22128 ) (30) The twists and eigencompliances are ordered by decreasing compliance. The compliance matrix can be obtained using Equation 8. The mechanism is actuated around the first eigentwist; the point where the leaf springs cross. The desired response is a parallel moment. Since the applied twist is an eigentwist, this is indeed the case. When the mechanism is deflected as seen in Figure 4, the induced twists, eigentwists and eigencompliance change. In a deflected state of 30 degrees, they are as follows. T\u0302 = \u22120.0227 \u22120.0086 \u22120.0103 0.0039 0.5199 \u22120.8542 0.0014 0.0312 \u22120.0137 0.9998 \u22120.0184 \u22120.0066 0.0103 0.0010 0.0022 \u22120.0192 \u22120.8540 \u22120.5199 0.0196 0.5200 \u22120.8540 0 0 0 0.9988 \u22120.0492 \u22120.0071 0 0 0 \u22120.0457 \u22120.8528 \u22120.5203 0 0 0 (31) c\u03b5 = diag(0.42 0.051 0.017 2.66\u00b710\u22127 7.83\u00b710\u22126 1.16\u00b710\u22125 ) (32) If the original twist is now applied, the response will be different. The linear combination is obtained using Equation 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003100_j.ifacol.2018.11.569-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003100_j.ifacol.2018.11.569-Figure5-1.png", + "caption": "Fig. 5. Geometrical model of a safety zones around an obstacle, qualitative illustration.", + "texts": [ + "1 the CoM of a human person cannot contact the obstacle\u2019s boundary. A person maintains a minimum distance to the obstacle due to the width of the human body, and some additional distance to reduce the risk of collision. A potential field component is often used in the cost function to represent this distance around the obstacle. However, for reasons already discussed in section 3.2, F is chosen as a simple quadratic equation of the position deviations in this work. Instead, we propose a geometrical model of a safety zone around the obstacle (Fig. 5). This geometrical model is given to the NMPC as constraint on the position components (x,y)lim. Through the usage of this safety IFAC SYROCO 2018 Budapest, Hungary, August 27-30, 2018 Alexander Joos et al. / IFAC PapersOnLine 51-22 (2018) 366\u2013371 369 NMPC, as well as limitations for the walking speed V. This limitation on the forward speed is used to force the NMPC to choose the same \u201cwalking\u201d speed as the human participants in the experiments. The rotational velocity \u03c9, which is the rate of change of walking direction \u03c8, is not limited", + " (2016)), we will show in this document, that the CoM trajectories can be approximated accurately through a suited choice of the prediction horizon \ud835\udc47\ud835\udc47\ud835\udc5d\ud835\udc5d with this cost function. As common for NMPC, the optimal input is applied over a time interval \u03b4 (NMPC update or sample time) as shown in Fig. 4 (grey field). After time interval \u03b4, which is significantly shorter than the prediction horizon \ud835\udc47\ud835\udc47\ud835\udc5d\ud835\udc5d in this work, the optimization (Eq. (2)) is restarted from the new actual time \ud835\udc61\ud835\udc610 to calculate a new optimal input trajectory. Fig. 4. Standard NMPC principle. Fig. 5. Geometrical model of a safety zones around an obstacle, qualitative illustration. It is obvious that in the scenario shown in Fig.1 the CoM of a human person cannot contact the obstacle\u2019s boundary. A person maintains a minimum distance to the obstacle due to the width of the human body, and some additional distance to reduce the risk of collision. A potential field component is often used in the cost function to represent this distance around the obstacle. However, for reasons already discussed in section 3.2, F is chosen as a simple quadratic equation of the position deviations in this work. Instead, we propose a geometrical model of a safety zone around the obstacle (Fig. 5). This geometrical model is given to the NMPC as constraint on the position components (x,y)lim. Through the usage of this safety 0t 0t pTt 0 statereference predicted x predictedu 0x 0u IFAC SYROCO 2018 Budapest, Hungary, August 27-30, 2018 zone model, the NMPC can imitate the human CoM trajectories around the obstacle with a high accuracy as will be shown below. The center of the safety zone is shifted in x-direction relative to the center of the obstacle by \ud835\udc51\ud835\udc51\ud835\udc65\ud835\udc650, perpendicular to the main direction of motion", + " The angle \u03b1 gives their relation: tan \ud835\udefc\ud835\udefc = \ud835\udc51\ud835\udc51\ud835\udc66\ud835\udc660 \ud835\udc51\ud835\udc51\ud835\udc65\ud835\udc650\u2044 . The distance between maximal expanse of the obstacle and the safety zone in x-direction is denoted with \ud835\udc51\ud835\udc51\ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a. For small \u03b1, which is typical for the models found from experiments, \ud835\udc51\ud835\udc51\ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a\ud835\udc5a is a good approximation of the minimal space between the human CoM trajectory and the obstacle. This minimal space between the human and the obstacle is a geometrical feature that will later be used for the systematic description of the generalized safety zone model. The illustration in Fig. 5 is valid in a scenario where the human walks around the obstacle on the obstacle\u2019s left side while walking in positive y-direction. If the human walks on the right side of the obstacle, the safety zone model is mirrored at the line x = 0m respectively. 4. MODEL PARAMETERS AND VALIDATION The numerical values of NMPC parameters \ud835\udc47\ud835\udc47\ud835\udc5d\ud835\udc5d and \u03b4 (Joos and Fichter (2018)), as well as the geometry of the safety zone model have a crucial impact on the geometry of the resulting position trajectory. Therefore, suited parameters of the NMPC and the safety zone model need to be developed to enable an accurate approximation of the human CoM trajectories" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002962_icmic.2018.8529920-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002962_icmic.2018.8529920-Figure8-1.png", + "caption": "Fig. 8 Different faults of the rotor-bearing system By using the NOFRFs, the results of 2 ( j2 ) F", + "texts": [ + "6 is tested in the y direction, where, in order to evaluated the NOFRFs of the system under different working conditions, the imbalance distance e is adjusted by changing the position of two additional bolts with the mass of 39.7 10 kgm as shown in Fig.7, and the equivalent imbalance distance e are listed in Tab.III. TAB.III EQUIVALENT IMBALANCE DISTANCE i e i imbalance distance/ mm 1 110.44 2 79.62 3 21.97 The test was conducted under the rotating speed of 2000 r/ min r n , where the input of the rotor-bearing system can be formulated as 2 sin F F tu me t (21) where 209.44 rad/s F . The faults of bearings are simulated by putting different objects under the bearing seat as shown in Fig.8, where different layers of A4 papers and metal shims are applied in the test. G under different faults are shown in Fig.9, comparing with the detection results by only using ( j ) F Y under the imbalance distance 1e in Fig.10. In Figs. 9 and 10, Sn and Pn represent n layers of shims and papers, respectively. VI. CONCLUSIONS The NOFRFs of nonlinear systems are the natural extension of the FRF for linear systems, which can be used for nonlinear systems analyses and design in a wide range of practical applications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002089_iscaie.2018.8405443-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002089_iscaie.2018.8405443-Figure2-1.png", + "caption": "Fig. 2. Side view of the machine.", + "texts": [ + " The corresponding x, y, and z components of the ball\u2019s translational and angular velocity were represented by (Vx ) , (Vy ) , (Vz ), \u03c9x, \u03c9y, and \u03c9z respectively. max= - 1 2 CD\u03c1A V .Vx- 1 2 CL\u03c1A V Vz\u03c9y-Vy\u03c9z \u03c9 (1) may= - 1 2 CD\u03c1A V .Vy- 1 2 CL\u03c1A V Vx\u03c9z-Vz\u03c9x \u03c9 (2) maz= - 1 2 CD\u03c1A V .Vz1 2 CL\u03c1A V Vy\u03c9x-Vx\u03c9y \u03c9 (3) V= (V1+V2) 2 (4) CL=3.19\u00d710-1[1- exp -2.48\u00d710-3\u03c9 ] (5) II. METHODOLOGY The performance analysis of the cricket bowling machine was divided into three test groups namely velocity, distance, and spin rate tests. The CAD of the fabricated machine used for the tests is illustrated in Fig. 1 and Fig. 2. 978-1-5386-3527-8/18/$31.00 \u00a92018 IEEE 47 A. Velocity test The velocity experiments were done on the machine to investigate the effects of two manipulated variables namely ball wetness and wheel pressure on the ball\u2019s translational velocity. The device used to measure the ball\u2019s initial velocity was a camera with high speed recording at 1000 frames per second placed at a distance on one side of the machine. On the opposite side, a black board with two vertical white lines was placed as the measuring scale", + " The neutral line referred to the trajectory with no ball spin. It can be seen in Fig. 6(c) that under Rule 11, the ball went further to the right but under Rule 20, it drifted in the opposite direction. This was also similar to Rule 10 and 19 which had the ball launched to left side. Rule 10 applied top and left spin whereas Rule 19 applied back and right spin. IV. DISCUSSION The speed ratio of the ball\u2019s linear speed to that of the wheel differs between the dry and wet ball tests as shown in Figure 2. Both test showed that the speed ratio is inversely proportional to wheel speed which coincides with past experimentation results on similar machines across varying wheel speed range as shown in Fig. 9 [19]. Even at large wheel angular speed, the speed ratio maintained a relatively linear relationship. However, the gradient magnitude of past machines\u2019 speed ratio plot is smaller compared to that of the current two-wheel machine. The 900 rpm data for dry and wet ball were added to obtain a better approximation of the best fit line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003311_smc.2018.00076-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003311_smc.2018.00076-Figure3-1.png", + "caption": "Figure 3 Vibrator", + "texts": [ + " Goniometers were attached to participant\u2019s both elbow joints. By using the potentiometers (ALPS, RDC501051A) of these goniometers, the participant\u2019s both elbow joint angles were measured. The signals from these potentiometers were conveyed to a microcontroller (Arduino UNO) . The microcontroller converted the analog signals from the potentiometers into the digital signals. The digital signals were finally conveyed to a laptop PC from microcontroller for recording the measured data. A connection diagram of the experimental device is shown in Fig. 3. The vibrator is shown in Fig. 4. This vibrator can generate vibration stimulation with piston-crank mechanism. A DC motor (Mabuchi motor, RS380-PH) was used in this vibrator. As shown in Fig. 4(b), the rotation shaft which was attached to the DC motor has 1mm eccentricity. This eccentric plays the role of crank and defines the amplitude of vibration stimulation 1 (mm). The frequency of vibration stimulation was changed in each trials of the experiments. This vibrator can generate 20 to 90 (Hz) vibration stimulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000415_1.1718791-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000415_1.1718791-Figure1-1.png", + "caption": "FIG. 1. Electrode assembly for anodization of semiconductors.", + "texts": [ + " This article describes such an oscillator capable of delivering 20 W at 115 V ac at any arbitrary frequency in the range from 25 cps to 12 kc. It will match correctly a variety of impedances, has good waveform when ~687 c ~t- '''~''WO'' , , I 02 I \" \"I \",.1\"'1 rF OSCILLATOR VOLTAGE FEED-BACK CONTROL ADJUSTMENT connected to a motor load, has a high degree of frequency stability, has provision for intermittent on-off or remote operation, and has a finely adjustable frequency (i.e., speed) control. The first stage (Fig. 1) is a \\Vien bridge oscillator employing a 6CG7 double triode. The frequency delivered by the bridge is given by the relation frequency= 1/ (27rR tC), where C is the capacitance of a bridge ann, and R t is the total resistance in a bridge arm. In this case, R t is the sum of the fixed resistance Rl and the variable resistance Rv The purpose of the fixed resistance in the bridge arms is to compress the range of variation and thus provide finer frequency control. If desired, several ranges can be in cluded by providing the appropriate switching", + " It is desirable to use choke input since keying the cathode of the power stage removes the major portion of the power-supply load that, with a capacitor-input supply, would cause the voltage to soar nearly to the peak trans former voltage and thus require the use of 700-V filter capacitors, which are bulky and expensive. General construction is not critical, and the use of triodes provides a degree of stability not readily obtainable with pentodes. In line with good construction practice to prevent unwanted feedback, parts layout should follow the general geometrical arrangement of the schematic. 6550 II~ FIG. 1. Schematic of variable-frequency power oscillator. All capacitances are in.uF. This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitationnew.aip.org/termsconditions. Downloaded to IP: 130.209.6.50 On: Thu, 18 Dec 2014 21:59:37 The only pre-use adjustment necessary is that of the oscillator feedback control. This control is adjusted for least distorted output waveform, or in the absence of distortion-analysis equipment, minimum amplitude out put at which stable oscillation still exists" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003949_iscid.2018.10152-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003949_iscid.2018.10152-Figure2-1.png", + "caption": "Figure 2. Meshing diagram of the driving drum", + "texts": [ + " In this step, the contact force between the rigid shaft and the inner wall of the drum is simulated and analyzed. The definition time is 1s. The second analysis step is defined as the pressure analysis step, which simulates the stress and strain analysis of the belt pressing the drum through the drum, and the time is also defined as 1s. At the same time, a monitoring point is set up. In order to study the effect of stiffening ring on the overall stress and strain of the drum, the highest point RF of the stiffening ring corresponding to the center position of the drum wall shown in Figure 2 is selected to monitor the change of the degree of the X-axis direction in each analysis step increment. There is no need for mesh generation for axis because it is defined as a rigid body. For the driving drum, a 4-node linear tetrahedron element C3D4 is used to mesh the finite element method. A total of 8201 nodes and 25143 elements are used. The meshing diagram is shown in Figure 2. As mentioned earlier, if the shaft and expansion sleeve are regarded as a rigid body, the constraint of the rigid body on the cylinder is the only boundary condition of the model. Defining the boundary conditions requires defining each analysis step. In the contact analysis step, all degrees of freedom of the rigid body are restrained. In the pressure analysis step, the UR2 degree of freedom is relieved, so that the rigid body can rotate around the Y axis. At the same time, the boundary condition of monitoring point RF is defined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003050_jae-171190-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003050_jae-171190-Figure2-1.png", + "caption": "Fig. 2. 3D-CAD model of the electromagnetic.", + "texts": [ + " Therefore, reduction in size of the entire apparatus is expected. It is also expected to reduce power consumption. In this research, we develop linear actuator with features suitable for artificial hearts with such advantages. In addition, as a preliminary step to the application to the artificial heart, we applied it to the electromagnetic pump and decided to evaluate the basic structure and characteristics of the experiment. Figure\u00a01 shows the pulsating electromagnetic pumps developed using permanent magnet linear actuator. Figure\u00a02 shows 3D-CAD model of the electromagnetic pump. The electromagnetic pump newly developed in this research is mainly composed of permanent magnet, tank, driving part. The drive uses a DC motor, and the reduction ratio between the motor and the permanent magnet is 1:180. We use 4 poles, 6 poles, 8 poles for permanent magnets. Each magnet reciprocates 90 degrees for 4 poles, 60 degrees for 6 poles and 45 degrees for 8 poles. The electromagnetic pump performs one beat by reciprocating the magnets once" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002408_978-3-319-99522-9_9-Figure9.9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002408_978-3-319-99522-9_9-Figure9.9-1.png", + "caption": "Fig. 9.9 Kinematical scheme of leg A of the mechanism", + "texts": [ + " In that following, we focus on an efficient and fast recursive matrix method, which is adopted to derive the kinematic model and the inverse dynamics equations of a three-degree-of-freedom symmetric spherical parallel robot with revolute actuators and three identical legs. The spherical manipulator has a construction with the axes of all nine revolute joints concurring in a common center O of rotation of the device. The mechanism input of the robot is structured into three active revolute joints while the output body is connected through a set of legs with identical topology [20]. Let Ox0y0z0\u00f0T0\u00de be a fixed Cartesian frame, about which the manipulator moves. It has three legs of known size and mass (Fig. 9.9). Each of the legs of this robot is a serial 3-RRR symmetric spherical wrist: the first joint couples the base with the proximal link, the second joint couples the proximal link with the distal link, and the last one couples the distal link with the end-effector (Fig. 9.8). The first element T1 of leg A(aA \u00bc 0), one of the three driving components of the robot, is called the proximal link. It is a homogenous rod, rotating about the axis OzA1 with the angular velocity xA 10 \u00bc \u20acuA 10 and the angular acceleration eA10 \u00bc \u20acuA 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003358_phm-chongqing.2018.00035-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003358_phm-chongqing.2018.00035-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of planetary gear structure", + "texts": [ + " It has been widely applied in aerospace, metallurgy, power, shipbuilding and other industries. Once the planetary gearbox fails, the equipment and the whole power transmission system will be destroyed, and the consequences are extremely serious [1]. Therefore, it is of great practical significance to carry out the research on the state monitoring and fault diagnosis of the planetary gearbox [2]. Planetary gearbox is generally composed of sun gear, planetary gear, ring gear and other key components. Fig.1 shows a schematic diagram of the planetary gear structure with three planetary wheels. As a complex mechanical system, when the planetary gear box fails, its dynamic behavior often appears complexity and non-linearity, and its vibration signal also presents non-stationary. At the same time, the periodic impact caused by the gear fault will play a role of amplitude modulation and frequency modulation for meshing vibration. Therefore, the fault vibration signal of the planetary gearbox is a non-linear and non-stationary signal with the characteristic of amplitude modulation and frequency modulation [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002337_icphm.2018.8448881-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002337_icphm.2018.8448881-Figure8-1.png", + "caption": "Figure 8. The experimental bearing", + "texts": [ + " The structure of the data acquisition and monitoring system of the experimental platform for monitoring the lubrication state of the rolling bearing under normal conditions is shown in Figure 7. The experimental sample contains 7 kinds of bearings with different grease contents. Each type contains 2 sample bearings and acquires 15 groups of AE and vibration signals at different speed. The total data collection is 480 groups, as showed in Table 1. The experimental bearing is the Rolls-Royce 6308 model of the rolling bearing, the bearing physical map shown in Figure 8. 978-1-5090-0382-2/16/$31.00 \u00a92018 IEEE During each rolling bearing experiment, the motor speed rose steadily from 100r/min to 1500r/min, and the AE and vibration acceleration signals were collected once every 100 revolutions and kept for at least 30s at each speed before the collection of data to ensure that each signal acquisition are carried out in a stable bearing operating conditions. IV. INFLUENCE OF GREASE ON LUBRICATION OF ROLLING BEARINGS Firstly, since the original signal collected in the experiment contains a lot of noise, which is not conducive to signal feature extraction, it is necessary to filter the original signal effectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000583_ichr.2008.4755949-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000583_ichr.2008.4755949-Figure6-1.png", + "caption": "Fig. 6 The kinematics of the PA-I 0 robot is illustrated. The wrist position inverse kinematics can be solved analytically when only the angles 8\\ 82 84 are considered and 83 is taken as zero. Note that the wrist position does not depend on 85 86 87 .", + "texts": [ + " Given a desired wrist position on the PA-I 0 arm, it is possible to analytically find the joint angles that would achieve the desired wrist position. Since the world coordinate frame is chosen to be aligned with the robot reference frame, one can directly use human wrist position data (in our world coordinate frame) for the inverse kinematics. When only PA-I0 arm joints 81, 82, 84 are considered the geometry allows the analytic calculation of the required angles for achieving the desired wrist position p (see Fig. 6). , camera camera 1) Pi == Pi - P origin ( p;vorld == WRF *P; (2) The hand reference frame is defined by the orthogonal unit vectors as illustrated in (see Fig.5). The z-axis is along the finger tips when hand is fully extended, and the y-axis is approximately aligned with the vector from index finger knuckle to little finger knuckle. And x-axis is the cross product ofy-axis and z-axis. 100 ..,.JOorMntItIon,....) no (c) oIIjtctlt.nlltlon (1Irw. no (a) 140 with uncommon task requirements (i.e. prying a nail versus hammering a nail) yield more options for grasping, and hence lead to an increased variance" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001631_0954406218763699-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001631_0954406218763699-Figure4-1.png", + "caption": "Figure 4. Application.", + "texts": [ + " Knowing the value Pf, from expression (35) v f i n o , i \u00bc 1, 2 is determined; with the aid of the second relation (19) Ef ic are calculated; using relation (42) Ef c is determined, and with the aid of the first and third relation (41), the coefficients of restitution kN and ke are calculated. The energetic variant From the third relation (41) Ef c is determined, and then equation (39) is integrated until the value given by expression (42) becomes equal to it. In this way, v f 12 n o and Pf are resulted. Further on, from the first equations (41) kN and kP are resulted. Application Collision between a homogeneous bar of constant transversal section (Figure 4) having mass m and length l, with a fixed plane knowing that v0x \u00bc v0, v0y \u00bc l1v0 cos 0, v0z \u00bc l1v0 sin 0, !0x \u00bc !0, !0y \u00bc l2!0, !0z \u00bc l3!0 is studied. Numerical application for l \u00bc 1 m, m \u00bc 1 kg, r \u00bc 0:006 m (radius of the circular transversal section of the bar), v0 \u00bc 10 m=s, l1 \u00bc 0, 5, 0 \u00bc 30 , \u00bc 45 , !0 \u00bc 5 rad=s, l2 \u00bc l3 \u00bc 2, \u00bc 0, 2 , and k \u00bc 0, 7, where k stands for kN, k , ke. Solution. In this situation, we may write v012x \u00bc v0x \u00bc 10 m=s, v012y \u00bc v0y \u00bc 4:330127 m=s, v012z \u00bc v0z \u00bc 2:5 m=s, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000012_6.2008-4509-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000012_6.2008-4509-Figure6-1.png", + "caption": "Figure 6- 360o Composite Ring Assembly", + "texts": [], + "surrounding_texts": [ + "American Institute of Aeronautics and Astronautics\n092407\n5\nThe test rig shown in Figures 2 and 3 has the following capability:\n\u2022 12,000 rpm \u2022 Delta Pressure Difference of 100 psid \u2022 Ambient Temperature Operation\nTesting was undertaken, using several seal configurations, rotational speeds, clearances and pressure\ncombinations on the following seals:\n\u2022 1thru 4 tooth labyrinth metal seals\n\u2022 1 split ring rotating brush seal\n\u2022 1 360 o ring rotating brush seal\n\u2022 2 split ring rotating brush seals arrange in tandem\n\u2022 2 360 o ring rotating brush seals arrange in tandem\n\u2022 2 split ring rotating brush seals arrange in tandem with centering springs\n\u2022 1 fractured carbon ring\nThe parts tested, with the exception of the carbon seals, are shown in Figures 4 thru 9.", + "American Institute of Aeronautics and Astronautics\n092407\n6", + "American Institute of Aeronautics and Astronautics\n092407\n7" + ] + }, + { + "image_filename": "designv11_92_0001813_978-981-10-8306-8_13-Figure13.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001813_978-981-10-8306-8_13-Figure13.1-1.png", + "caption": "Fig. 13.1 Schematic view of a the yield surface and b the effective plastic strain increment surface in the nine-dimensional space constructed by the plastic work equivalence principle", + "texts": [ + " Similarly, for the plastic strain increment, a conjugate quantity known as the effective (or equivalent) strain increment d e exists, whose expression of d e\u00f0dep\u00de \u00bc constant describes a surface as well as a generalized magnitude of the plastic strain increment. The effective plastic strain increment, which is a conjugate to a given effective stress, or vice versa, is defined by the following plastic work equivalence principle: dwp \u00bc tr\u00f0rdep\u00de \u00bc r dep \u00bc rijde p ij\u00f0\u00bc r1ijde 1;p ij \u00bc r2ijde 2;p ij \u00bc . . .\u00de \u00bc r\u00f0r\u00ded e\u00f0dep\u00de \u00bc constant \u00f013:4\u00de To develop the conjugate, there are two procedures to apply the plastic work equivalence principle: which can be described as figurative and algebraic procedures. As for the figurative procedure, consider Fig. 13.1, in which the stress and plastic strain increments are considered as nine-dimensional vectors and dwp is their dot product, while the stress stays on the yield surface whose size is defined by its reference stress: r\u00f0r\u00de \u00bc constant. As for the plastic strain increment, its direction is normal to the yield surface and, when the size of the dot product of the two vectors are fixed as dwp = constant, its size is determined by Eq. (13.4) so that the plastic strain increment vectors construct a surface, whose size is also defined by its reference plastic strain increment; i.e., d e\u00f0dep\u00de \u00bc constant \u00bc dwp=constant defined by r\u00f0r\u00de \u00f013:5\u00de which defines the effective plastic strain increment and its surface is known as the effective plastic strain increment surface. Furthermore, for the constant plastic work increment, Eq. (13.4) becomes d\u00f0dwp\u00de \u00bc dr dep \u00fe r d\u00f0dep\u00de \u00bc 0 \u00f013:6\u00de where dr is the increment between the neighboring stresses on the yield surface; therefore, dr is tangential to the yield surface as shown in Fig. 13.1a. Then, by the normality rule of Eq. (13.3), dr is normal to dep so that dr dep \u00bc 0. Consequently, r d\u00f0dep\u00de \u00bc 0, in which d\u00f0dep\u00de is tangential to the effective plastic strain increment surface and r is normal to the surface as shown in Fig. 13.1b; i.e., r \u00bc A @dg\u00f0dep\u00de @ dep\u00f0 \u00de \u00bc A @d e\u00f0dep\u00de @ dep\u00f0 \u00de \u00f0\u00bc r @d e\u00f0dep\u00de @ dep\u00f0 \u00de \u00de \u00f013:7\u00de where dg\u00f0dep\u00de is the plastic strain increment function. The plastic strain increment function defines the effective plastic strain increment surface just as the yield function defines the yield surface, and whose first order homogeneous function is the effective plastic strain increment. The two normality rules shown in Eqs. (13.3) and (13.7) are dual normality rules for the conjugate effective stress and the effective plastic strain rate increment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000418_ma901483d-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000418_ma901483d-Figure1-1.png", + "caption": "Figure 1. Deformation of the helix structure by tensile stress increasing from left to right. For the two rightmost structures, the deformations exceed a critical value, resulting in only oscillating directors, no longer showing complete rotations.", + "texts": [ + " A unique feature of cholesteric liquid crystals is the selective reflection of circularly polarized light with the same handedness as the helix in a narrow range around a center wavelength \u039bR,0. 2 The transmission of oppositely polarized light is not affected significantly. The spectra are very sensitive to helix distortions. For tensile stress applied perpendicular to the helix axis of cholesteric elastomer films and a deformation exceeding a critical value, untwisted states were predicted theoretically,3,4 where the director only shows oscillations (see Figure 1). However, experimental studies5 showed serious deviations from the theoretical expectations. The transmission of polarized light indicated a helix distortion much weaker than predicted. In particular, no clear indications for the existence of untwisted states were found. Later, the contribution of the Frank elasticity (the elastic response to the distortion of a liquid crystal\u2019s director field) to the free energy density, previously neglected in comparison to the network elasticity, was studied theoretically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002073_ccdc.2018.8407818-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002073_ccdc.2018.8407818-Figure2-1.png", + "caption": "Fig. 2 Optimal Advance Angle Solution", + "texts": [ + " with 1 2=\u03c8 \u03c8 The spatial distribution error of ATT and incoming torpedo obeys normal distribution During the ATT intercepts incoming torpedo and ATT has a wide search sector angle and homing distance, Based on the single peak and symmetry probability of normal distribution density, When the torpedo's limit search angle is equal, The angle between the main direction of the coming torpedo and the azimuth line of the target is the optimal advance angle, Limit angle refers to the angle between the ligature of 4002978-1-5386-1243-9/18/$31.00 c\u00a92018 IEEE target and torpedo and the ligature of target and ATT lateral point search sector the target can be found by ATT when the target motion within the limit angle. Taking the target to the right as an example, the optimal advance angle is calculated as follows: As showed in Figure 2, \u03bb is search sector angle, make 1 2\u03c8 \u03c8 \u03c8= = ,and a z\u03d5 \u03d5 \u03b3= \u2212 (2) In triangle MLF: sin( / 2 ) arctan cos( / 2 ) a a a D R \u03bb \u03d5\u03d5 \u03b3 \u03bb \u03d5 \u2212 \u2212 = \u2212 \u2212 (3) In triangle MEL: sin( / 2 ) arctan cos( / 2 ) a a a D R \u03bb \u03d5\u03d5 \u03b3 \u03bb \u03d5 + + = \u2212 + (4) It can be obtained by (3)(4): 1 2 sin( / 2 )arctan cos( / 2 ) sin( / 2 ) arctan cos( / 2 ) a a a a a a D R D R \u03bb \u03d5 \u03bb \u03d5 \u03b3 \u03bb \u03d5 \u03bb \u03d5 + \u2212 + = \u2212 \u2212 \u2212 \u2212 (5) sin sin( )m z z V V \u03d5 \u03b8 \u03d5= \u2212 , sinarctan 1 cosz m m \u03b8\u03d5 \u03b8 \u22c5= + \u22c5 ,in which mVm V = , a mQ\u03b8 \u03d5= + . The parameters in the upper form are as follows: tV \u2014Torpedo speed gV \u2014Target speed Qm \u2014The angle between the line of the submarine and the direction of the target velocity D \u2014Initial launch distance R a \u2014Torpedo self guided distance \u03bb \u2014The fan angle of torpedo Unlike naval vessels, the ATT is launched at the 90 broadside of the submarine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002240_msf.928.209-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002240_msf.928.209-Figure1-1.png", + "caption": "Fig. 1. Structure of bimorph thermal actuator", + "texts": [ + " When one of the bi-layer beam ends is fixed, a specific displacement to be measured at the other end. The high CTE material exhibits compressive stress and low CTE material exhibits a tensile stress. Structure of Bimorph. The polymer composite materials bonded at an interface for making bimorph beam. The two layers made up of pure PDMS and CB-PDMS composite polymer. The bilayer has the same length, width and varying thickness are represented as L, W, t1, t2 respectively. The bimorph structure proposed has 90 mm length, 14 mm width and 10 mm thickness as shown in Fig.1. Three thicknesses of bimorph between 4 mm to 6 mm are also studied. The model is simulated using multiphysics software. The reference temperature is used for calculation is 20 0C. The analysis is carried out for temperatures 20 0C to 80 0C in steps of ten degree Celsius. Required properties of materials for theoretical and simulation are shown in Table.1 Sample Coding. Two variables are considered for analysis, thickness and volume percentage of filler material. These combinations are represented using the codes shown in Table.2. The code BM605 indicates bimorph actuator of 6 mm thick CB-PDMS layer and 5 Vol% of carbon black filler. The rest of the paper uses this code for identification of thickness and filler ratio. Analytical Model of Proposed Bimorph. The thermal actuator consists of two layers, that influence its displacement. Parameters influencing the above are CTE, strain, Young\u2019s modulus and temperatures of both layers are shown in Fig.1 (a). The bimorph going in to difference in CTE takes a shape, when heated as shown in Fig.1 (b). The radius of the curvature (R) of this shape is given by Eq. below [11]. ( )ttxyzyzxyzxzyx T)\u03b1\u03b1(x)(xyz R 21)142634224( 1261 21 +++++ \u2206\u2212+ = (1) Where, ttx 12= ; wwy 12= ; EEz 12= ; EEttww and 212121 , are width, thickness and lateral Young\u2019s modulus of layer 1 and layer 2 of the polymer material respectively. )(, 21 \u03b1\u03b1\u03b1 \u2206 Is the CTE of their respective layers and T\u2206 is the temperature change between the initial and working temperature. Hence radius of curvature is proportional to the differences in CTE, temperature and inversely proportional to the thicknesses of the layers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003769_1.5092451-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003769_1.5092451-Figure4-1.png", + "caption": "Fig. 4. Schematic representation of the Alcon levitator.", + "texts": [ + " Placing a ring on the solenoid, concentric with the iron core and closing the circuit, produces a sudden current pulse on the ring, hence a force that throws the ring a certain distance up. Another possibility is to switch the current on and then place the ring on the solenoid. In this case, the ring levitates continuously, the iron core preventing it from lateral displacements. Alcon\u2019s levitator3,12 makes use of the induced current on a moving conductor.13 A schematic representation of the geometric configuration is shown in Fig. 4: two cylindrical metallic pieces rotate at the same angular velocity and a magnet is placed above the cylinders in the middle plane. The net force that the induced currents exert on the magnet is upward. If the rotation of the cylinders is fast enough, the magnet can be levitated. Since levitation requires rather high velocities, we decided, for safety reasons, not to propose that our students attempt a levitation experiment. Instead, we assembled a setup to quantify the effect and, eventually, calculate the angular speed required for levitation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002887_ihmsc.2018.00016-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002887_ihmsc.2018.00016-Figure2-1.png", + "caption": "Figure 2. Schematic of deflected slipstream STOL[6].", + "texts": [ + " The strake is employed to improve the aerodynamic characteristics of the wing at a large angle of attack, as can been seen in Figure 1. 32 978-1-5386-5836-9/18/$31.00 \u00a92018 IEEE DOI 10.1109/IHMSC.2018.00016 The deflected slipstream to increase the lift is enabled by the double slit flaps which change the direction of the propellers slipstream. The resultant force of thrust, lift and drag is supposed to be vertical to offset part of gravity and thus achieving short takeoff and landing [6], as shown in Figure 2. The shafts of main propellers are parallel with the centerline of the fuselage. In order to get a vertical force for take-off and landing, a certain angle between the body and the ground shall be maintained. As proved by previous research, it is feasible to maintain this angle by installing a propeller vertically at the nose of to the aircraft. This arrangement enables the angle control and moment balance before take-off, as can be seen in Figure 3. III. MATHEMATICAL MODEL OF THE V/STOL This V/STOL aircraft involves a vertical tension for the balance of force, and the impact of the propeller deflected slipstream shall be considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001858_s12650-018-0495-1-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001858_s12650-018-0495-1-Figure2-1.png", + "caption": "Fig. 2 The coordinate system and nomenclature for the tumbler system are shown. The streamlines and boundary layer depth, d, are illustrated for flows that circulate: rolling, cascading and cataracting", + "texts": [ + " While some studies suggest that using colored, flagged or tracer particles might glean more information from a single run of particles within a three-dimensional field (Bendicks 2011), such methods have difficulty tracking all the particles through time. When these particles move in and out of the flow fields, tracking of individual particles is lost and complete time data histories are not feasible (Homeniuk et al. 2010). Instead, a pseudo-two-dimensional field was utilized in this study. The work presented here could be extended to three-dimensional flows where the tracking of individual particles would be possible as long as they remain in the field of view. 2 Experimental method 2.1 Setup Figure 2 presents a schematic of the geometry and nomenclature used. A 13.97 cm inside diameter tumbler, D, was constructed from clear acrylic and was driven by a variable speed brushless motor. A phototachometer was used to measure the clockwise (relative to the laboratory frame of reference) tumbler rotation rate, X, which was as high as - 5 revolutions per second (rps). The tumbler was mounted on soft, freely rotating, rubber casters with internal cylindrical roller bearings to provide smooth operation. The tumbler contained an interior wall that could be adjusted to accommodate either single row or multiple rows of particles for both two- and three-dimensional analysis. Dashed lines in Fig. 2 represent streamlines of particles traveling through an averaged flow field. These particles change direction, indicated with a solid black line and denoted as the shear layer thickness, d(x), where the boundary layer thickness, d, is the depth at x = 0 cm. The streamlines are valid for circulating flows, such as rolling, cascading and cataracting. Finally, Fig. 2 contains the representation of the dynamic angle of repose, b0 (Pignatel et al. 2012; Duong et al. 2004). The particles utilized in this study were right-circular cylinders. The diameter of each particle, d, was 1.25 cm with an axial length of 0.30 cm. Each particle weighed 0.42 g and was laser cut from a single white ACRYLITE acrylic sheet to insure consistency and accuracy of construction. This uniformity prevented the \u2018\u2018Brazil Nut Effect,\u2019\u2019 common to S- and D-systems, where larger particles rise to the surface of a grain bed (Rosato et al", + " The resulting modes obtained were qualitatively similar to those where the bed depth is much greater. 3.2 Translational velocity Using the methodology described in Sect. 2, the translational velocity for each particle in the tumbler was measured for each mode of tumbler operation listed in Table 2. The transverse and streamwise velocities were non-dimensionalized by the tangential velocity of the tumbler and plotted as a function of bed depth in Fig. 6. Streamwise, u, and transverse, v, velocities refer to the x-component and y-component of particle velocity, respectively (see Fig. 2 for the coordinate directions). The average velocity as a function of depth is indicated with a solid line, while the dashed line indicates one standard deviation from the average. Finally, each individual particle is assigned a shade so that particle behavior can be assessed as a function of bed depth. Within the slipping mode, the particles were nearly stationary with respect to the coordinate system and their transverse velocities were well below the linear velocity of the tumbler. The non-dimensionalized velocities formed a symmetric velocity distribution about zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003138_978-3-030-02964-7_2-Figure2.7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003138_978-3-030-02964-7_2-Figure2.7-1.png", + "caption": "Fig. 2.7 (a) Sequencing graph for a multiplexed bioassay [24] (S1, S2 are samples, R1, R2 are reagents, M1 \u223c M4 are mixing operations, I1 \u223c I8 are dispensing operations, and D1 \u223c D4 are detection operations). (b) Constructed reservoir graph GR", + "texts": [ + " If there is an operation that relies on two reservoirs, i.e., an operation with droplets dispensed from reservoirs i and j , there is an edge eij in GR between the two corresponding vertices vi and vj . In the constructed reservoir graph GR , each edge has an associated numerical value, called a weight. The weight wij of eij is determined by the number of operations that correspond to reservoirs i and j . An example of the constructed reservoir graph GR from a multiplexed bioassay is shown in Fig. 2.7. The weight of each edge is highlighted in red. Since the time for droplet routing can be reduced if two corresponding reservoirs are close to each other, the objective function can be expressed as in (2.13). min : i =j\u2211 1\u2264i,j\u2264n wijDij (2.13) where Dij is the Manhattan distance between reservoir i and j , i.e., Dij =\u2223\u2223xi \u2212 xj \u2223\u2223 + \u2223\u2223yi \u2212 yj \u2223\u2223. The priority for each unscheduled operation is dynamically determined by the priority controller. The priority controller consists of a priority generator and a priority updater, where the priority generator is used to generate the initial priority for each operation, and the priority updater is used to update the priority whenever an operation has been scheduled or finished" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000983_icelmach.2008.4800111-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000983_icelmach.2008.4800111-Figure4-1.png", + "caption": "Fig. 4. Model of PMSG (Surface Mounted type) for FEA of field distribution. A-phase axis of winding is aligned with d-axis of rotor. Model covers only one pole pitch of generator. Field distribution for no-load condition is shown.", + "texts": [ + " It is based on the possibility of calculation (after previous FEA of field distribution inside the machine) of an internal voltage Ei and internal power angle \u03b4i values, corresponding to the actual field distribution at each particular machine operating point, assuming that a modulus I1 and a phase angle \u03b3 of an armature current phasor I1 are known \u2013 see Fig. 3 and Fig. 5. A model of PMSG for FEA is prepared so that the A-phase axis of stator winding is exactly aligned with the d-axis of rotor \u2013 see Fig. 4. Additionally, we assume that the FEA model corresponds always to an instant of time t = 0, and that the instantaneous values of phase currents, which are the input variables for FEA, are described by equations: In such a case, appearing in (2) the phase angle \u03b2 of the stator current phasor I1, calculated with respect to the rotor d-axis: \u03b2 = \u03c0/2 \u2013 \u03b3, corresponds simultaneously to the space angle between a stator magnetomotive force (MMF) vector Fs and a rotor MMF vector Ff (Fig. 5). Thanks to that the exact positioning of stator MMF in relation to d-axis of rotor in the FEA model is very easy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000021_imece2009-11165-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000021_imece2009-11165-Figure2-1.png", + "caption": "FIG. 2: MOVING INFINITE BAND HEAT SOURCE", + "texts": [ + " Additionally, normalized length coordinates are introduced: , 2 , 2 , 2 , 2 , 2 \u03ba\u03ba \u03ba\u03ba\u03ba lvLbvB zvZyvYxvX == === (8) whereas the parameter L is known as the P\u00e9clet number. This number states the ratio between heat conduction by convection and heat conduction by diffusion. For large P\u00e9clet numbers L, convective heat conduction prevails. It can be determined from this theory, that the greatest temperatures within the contact area occur along the center line which corresponds with the x-axis. If only the temperature along the centerline is looked at, a model of a band source, see Fig. 2, can be used. The band heat source moves into x-direction across the surface z = 0 of an infinite half space. Its length is 2l, while its width is infinite. Again, the velocity of the heat source is denoted by v. This center line temperature distribution can be calculated by Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 3 Copyright \u00a9 2009 by ASME \u03be\u03be \u03c0\u03c1 \u03be dZK cv qT LX LXp 22 0e +=\u0394 \u222b + \u2212 \u2212 (9) with the Bessel function of the second kind and zero order, K0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001637_j.proeng.2018.02.047-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001637_j.proeng.2018.02.047-Figure1-1.png", + "caption": "Figure 1 : hollow shaft and pinion assembly \u2013 main dimensions", + "texts": [ + "/ Procedia Engineering 00 (2017) 000\u2013000 In order to connect the electric motor to the gearbox without intermediate coupling device, Leroy Somer has developed the so-called \u201cMontage Int\u00e9gr\u00e9\u201d (MI, i.e. compact mount), where the input pinion is directly fixed on the motor shaft. This design requires a special shaft end. In the case of small pinions, the pinion shaft must be inserted in a hollow motor shaft end, with an additional feature for transmitting the torque: either a transverse pin, or a special key, as shown in Fig. 1. With the evolution of electric motors over the years, and specifically with the introduction of the IE2 efficiency compliant motor range, the torque characteristics of the motors have changed. For the most heavily loaded connections, it is therefore necessary to assess their mechanical strength. The objective of this work is threefold : 1. Detailed strength assessment : o What are the failure modes and where are the critical areas? o What is the allowable torque? This is not quite the same as getting a safety margin for a given load, it is actually the inverse problem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003102_ecce.2018.8558121-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003102_ecce.2018.8558121-Figure2-1.png", + "caption": "Fig. 2. Surface permanent magnet motor structure in xy cross-sectional view.", + "texts": [ + " PROPOSED MACHINE STRUCTURE Figure 1 shows a cut view of the fabricated prototype machine of the proposed single-drive bearingless motor by the authors for cooling fan applications [14]. This machine has a single-drive bearingless motor and two sets of repulsive passive magnetic bearings. At the top of the rotor shaft, a fan blade is installed. At the bottom, a displacement sensor is installed to detect the rotor axial displacement. One set of three-phase winding is wound in the center stator core as the concentrated winding. Figure 2 shows xy cross-sectional view of the single-drive bearingless motor. The cross section is identical with a surface permanent magnet machine so that the rotational torque is 978-1-4799-7312-5/18/$31.00 \u00a92018 IEEE 4392 generated by q-axis current. The numbers of rotor poles and stator slots are eight and twelve, respectively. The active axial force is generated by d-axis current in the single-drive bearingless motor. The detailed principle of the active axial force generation is shown in [14]. Figures 3 (a) and (b) show principles of passive stabilization in the radial and tilting motions by the repulsive passive magnetic bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002098_978-3-319-91590-6_6-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002098_978-3-319-91590-6_6-Figure4-1.png", + "caption": "Fig. 4 V-REP component reference axis details", + "texts": [ + " The image will be associated with a texture on a plane shape. To add the plane shape to the scene, click on Add \u2013> Primitive shape \u2013> Plane Set the plane dimensions to 10 \u00d7 10 cm. After that, select the plane object on the Scene Hierarchy menu and click on the Quick Texture button to choose the tag as a texture for the plane. This step finishes the composition of the workspace and its elements. Figure 3 shows the composed scene finished. The reference axis details of each scene component is shown on Fig. 4. The proposed application uses position data published to the AR-TAG ROS node to provide control information to a C++ code running on the simulation PC. This code performs all the control actions to lead the BOB robot from an initial position to a desired final set point relative to the quadcopter frame. All of BOB\u2019s moves will follow the quadcopter, so when the first changes its position, the other will move to achieve the previously set position defined by the code. The proposed architecture is composed of five elements on the simulation: 1 Quadcopter 2 Global Camera 3 AR -Track_Alvar Package 4 PID Control Script 5 BOB ground robot The quadcopter acts only as a transporter to move the Global Camera within the environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003949_iscid.2018.10152-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003949_iscid.2018.10152-Figure3-1.png", + "caption": "Figure 3. Load simulation diagram of driving drum", + "texts": [ + " The load of the roller shell is uniformly distributed from the conveyor. In this simulation analysis, the static dynamic analysis can simplify and omit the frictional force of the conveyor belt on the surface of the cylinder shell and the inertial force produced by the angular velocity of the cylinder in the running process. The pre tightening force of the expansion sleeve on the hub can be achieved through the contact of the rigid body to the drum. Therefore, the key point of the model is to realize load simulation, as shown in Figure 3. The circumference angle is 210 degrees. The load is uniform and the direction is perpendicular to the cylinder section. III. RESULT ANALYSIS The equivalent stress nephogram and the equivalent strain nephogram can be obtained by calculating the two steps with the finite element analysis software Abaqus. Figure 4 (a) is an equivalent stress nephogram of the drum under the load amplification of 500 times. For convenience of observation, the drum is cut symmetrically along the XOY plane to show the variation of stress inside the drum, as shown in Figure 4 (b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003671_uvs.2019.8658292-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003671_uvs.2019.8658292-Figure1-1.png", + "caption": "Fig. 1: Quadrotor configuration", + "texts": [ + " The section will present the application of GDI-SMC methodology for motion control of a rigid and symmetrical quadrotor system. To analysis the controller performance, the adopted system is X4-flyer quadrotor, whose major specifications are listed in Table I. Two reference frames are established, i.e., an inertial earthfixed reference frame denoted by E(xe, ye, ze), and the bodyfixed reference frame denoted by B(xb, yb, zb), whose center is coincided with the quadrotor\u2019s center of gravity as shown in Fig. 1. The inertial positions of quadrotor are represented by a vector \u03be = [xe ye ze] T , whereas the three Euler angles are denoted by the vector \u03b7 = [\u03c6 \u03b8 \u03c8]T bounded by \u2212\u03c0/2 < \u03c6 < \u03c0/2, \u2212\u03c0/2 < \u03b8 < \u03c0/2 and \u2212\u03c0 < \u03c8 < \u03c0 respectively. In Fig. 1, the symbols Ti, where i = 1, 2, 3, 4 represents the thrust force generated by the individual rotor and is calculated by Ti = b\u03c92 i (23) where b is the lift coefficient, and \u03c9i denotes the rotors angular speed. Similarly the control torques produced by the rotors along xb, yb and zb axis are given as \u03c4x = db(\u03c92 4 \u2212 \u03c92 2) (24) \u03c4y = db(\u03c92 3 \u2212 \u03c92 1) (25) \u03c4z = k(\u03c92 1 \u2212 \u03c92 2 + \u03c92 3 \u2212 \u03c92 4) (26) where d stands for the moment arm and k represents the dragto-moment coefficient, determined experimentally, [20]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001101_iros.2008.4651186-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001101_iros.2008.4651186-Figure9-1.png", + "caption": "Fig. 9. Overview of the experimental setup", + "texts": [ + " 8, it is expected that the rover having more than 12 [deg] of steering angles can traverse and climb the corresponding slope since the characteristics curve of \u03b2 = 12 [deg] is larger than the required cornering force and is included in the traversable region. A computational time for the above analysis takes around 0.5 [sec] (1.66 [GHz] processor). Most of the computational time spends to get whole data which are needed to form the thrust-cornering characteristic diagram. Slope traversal experiments using a four-wheel test bed were carried out to validate the proposed diagram for the trafficability analysis. In this paper, two cases are addressed; slope traversable and untraversable cases with changing slope angle. 1) Experimental Setup: Fig. 9 shows the overview of the experimental setup with the rover test bed. The test field consists of a rectangular tiltable vessel in the size of 2.0 by 1.0 [m]. The vessel is filled up with 8.0 [cm] depth of Toyoura Sand. The four-wheeled rover test bed has a dimension of 0.44 [m] (wheelbase) \u00d7 0.30 [m] (tread) \u00d7 0.30 [m] (height) and weights about 13.5 [kg] in total. Each wheel of the rover is same as the one used in the single wheel experiment. In the experiment, each steering angle of all wheels of the rover is controlled to maintain 15 [deg] to the uphill direction of the slope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002383_978-981-10-8597-0_9-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002383_978-981-10-8597-0_9-Figure1-1.png", + "caption": "Fig. 1 Schematics of the five-bar mechanism in different conditions", + "texts": [ + " Hence, the rank deficiency of J\u03b7\u03c6 does not necessarily result in the rank deficiency of J\u03b7q . The converse occurs when J\u03b7q becomes singular due to linear dependence of any two rows. Since, each row of J\u03b7q is a composition of a row of J\u03b7\u03b8 and J\u03b7\u03c6 by construction, this necessarily means, that the corresponding rows in the respective matrices also become dependent. Hence, the rank deficiency of J\u03b7q always implies rank deficiency of J\u03b7\u03c6 . To illustrate the above proposition, a planar five-bar mechanism has been used as an example (see Fig. 1). For this mechanism, \u03b8 = [\u03b81, \u03b82] , \u03c6 = [\u03c61, \u03c62] , q = [\u03b81, \u03b82, \u03c61, \u03c62] . Hence, m = n = 2 in this case. The constraint equations can be written as: \u03b71 = l cos \u03b81 + r cos\u03c61 \u2212 l0 \u2212 l cos \u03b82 \u2212 r cos\u03c62 = 0, (8) \u03b72 = l sin \u03b81 + r sin \u03c61 \u2212 l sin \u03b82 \u2212 r sin \u03c62 = 0. (9) The matrices J\u03b7\u03c6 and J\u03b7q are computed as J\u03b7\u03c6 = r (\u2212 sin \u03c61 sin \u03c62 cos\u03c61 \u2212 cos\u03c62 ) , J\u03b7q = r (\u2212\u03c1 sin \u03b81 \u03c1 sin \u03b82 \u2212 sin \u03c61 sin \u03c62 \u03c1 cos \u03b81 \u2212\u03c1 cos \u03b82 cos\u03c61 \u2212 cos\u03c62 ) , where \u03c1 = l r , is a nondimensional constant defining the ratio of the active to the passive link length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002497_978-3-319-99620-2_12-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002497_978-3-319-99620-2_12-Figure4-1.png", + "caption": "Fig. 4. Equivalent strain distribution on vessel for internal pressure of 1.5 MPa, (a) Strain near nozzle 1 is between 0.18 and 0.22%, (b) Max strain on inner surface", + "texts": [ + " 3). Applied boundary conditions and loads matched those used in experiments, while several different mesh densities had been used until the most appropriate was obtained. Final mesh consisted of 145,701 nodes and 72,732 solid elements. After applying pressure load, static nonlinear FEAs were carried out using material data obtained previously in the experiment with specimen [28]. The main purpose of these numerical simulations was verification of the developed model and \u2013 as it can be seen in Fig. 4 \u2013 strain values obtained in simulation were close to those obtained in the experiment. The highest strain values on outer surface appeared near nozzle 1 (Fig. 4a), in the same area where it appeared in experiment, with the values between 0.18% and 0.22% (max. value in experiment was 0.20%). Maximum strain in simulation 0.40% was observed on inner surface of a cylinder near nozzle 1 (Fig. 4b); however, this value couldn\u2019t be verified because no sensors have been used inside the pressure vessel. Nevertheless, values of strain measured on visible surfaces proved that numerical model was well defined. (It is worth mentioning that strain values near head\u2019s weld joint were about 0.15% and that maximum equivalent stress on vessel was around 200 MPa). After numerical model verification, next step in a research was optimization of the pressure vessel geometry with an aim of reducing its weight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001707_2018-01-0931-Figure15-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001707_2018-01-0931-Figure15-1.png", + "caption": "FIGURE 15 Rocker arm contact face diameter variations for impact study", + "texts": [ + " Minor drop in entrainment velocity is observed with increased base circle diameter as shown in Figure 13, with increased base circle diameter central oil film thickness at cam nose is increased slightly. Duration of critical minimum oil film thickness at zero/ minimum entrainment velocity is slightly changed as shown in zone A of Figure 14 with increased base circle diameter. Cam base circle has no major impact on cam shaft lubrication while it does reduce maximum cam contact stress. Impact of Rocker Arm Contact Face Diameter (for Sliding Contact) In sliding contact valve train mechanism rocker arm contact face diameter is varied as shown in Figure 15 which is in contact with cam profile due to its importance in cam contact tribology. Effective radius of \u00a9 2018 SAE International. All Rights Reserved. curvature at cam contact point is calculated using rocker arm contact face diameter and cam radius of curvature. Rocker arm contact face diameter is varied by 4\u00a0mm and 8\u00a0mm, with increase in diameter cam radius of curvature reduced at cam nose and increase at opening /closing flank. Increased contact face diameter increase contact stress at cam nose and reduce contact stress at opening /closing flank as shown in Figure 16" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001649_012093-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001649_012093-Figure2-1.png", + "caption": "Figure 2. Motion synthesis along the path prescribed by the segments \u0410s\u0410P1 and \u0410P1\u0410P2: a \u2014 with the deadlock occurrence at \u03b4 < 10 mm, b \u2014 motion synthesis at \u03b4 < 60 mm", + "texts": [ + " This component of the knowledge base characterizes the intrinsic properties of the android arm mechanism. Data on the start and end positions of points specifying the synthesized path of the OL motion and the forbidden region positions (of the tool racks P1 and P2) arise as the second component of the knowledge base 2, this knowledge base being the cause of deadlocks [10]. The example of the forbidden regions position P1, P2 and the path of the android robot arm motion, being the cause of the deadlock, is shown in figure 2a. The motion synthesis along the path \u0410S\u0410P1\u0410P2 with the point-to-point accuracy being \u03b4 < 10 mm is presented in figure 2a. In that case there is a deadlock. Figure 2b shows the motion synthesis along the same path with \u03b4 < 60 mm. 4 1234567890 \u2018\u2019\u201c\u201d Figure 2b shows that in the second case a significant deviation from the path occurs. Generally this can lead to the collision of the manipulation object and tool racks. In this case under the prescribed type of configuration, the motion synthesis in the vector of velocities for \u03b4 < 60 mm at the path segment \u0410P1\u0410P2 is impossible. Hence, providing the configuration at the point \u0410P1, first it is necessary to move the arm from the point \u0410S to the point AP1 with the type of configuration to be changed using the values of weighting factors for generalized velocity changes not being equal to unity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003639_jsuin.18.00059-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003639_jsuin.18.00059-Figure3-1.png", + "caption": "Figure 3. A freestanding smectic film two molecular layers thick with an island five molecular layers thick. The island has a disc-like form; in a freestanding film, the island diameter is much larger than its thickness. (a) Side view of the film; (b) top view of the film", + "texts": [ + " In the case of a defect on the boundary, the authors consider the contour around the defect between two points on the boundary on two sides of the defect. Figures 2(a) and 2(b) show the orientation of the c-director field near topological defects with topological charges S = +1 and S = \u22121. In the present work, the authors investigated orientational structures of the c-director field in smectic islands of antiferroelectric SmC*A liquid crystals.23,24 Smectic islands are disc-like regions of smectic films with the number of layers larger than that in the film (Figure 3). For rigid boundary conditions for the c-director on the boundary between the island and the film (e.g. the c-director is parallel or perpendicular to the boundary25), the total topological charge of defects inside the island and on the inner island boundary is fixed: SSi = +1.26 Despite a simpler geometry and geometrical restrictions related to the vector nature of the c-director, the possible number of combinations of topological defects in a smectic island can be larger than on a spherical surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001259_1.3183728-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001259_1.3183728-Figure1-1.png", + "caption": "Figure 1. Scheme of the planar sensor, with respective electrodes.", + "texts": [ + " The amperometric electrochemical sensors are devices that record electric current of a given electrochemical reaction at a constant potential. It allows one to qualify and to quantify substances through the relation between current response and their concentration (10). The electrochemical planar sensors were fabricated with three electrodes: working, reference and auxiliary, is described below. Sensor project: The fabricated sensor (1) is 17.0 mm width and 43.0 mm length with a working electrode of 2.5 mm in diameter, having alumina as substrate (Figure 1). The geometric configuration of the planar sensor based on a set of three electrodes was chosen to reduce the reference electrode area in order to permit a stable and fixed reference. 264 ) unless CC License in place (see abstract).\u00a0 ecsdl.org/site/terms_use address. Redistribution subject to ECS terms of use (see 128.255.-148.31Downloaded on 2015-06-01 to IP The sensor was defined by silk-screen technique: I) DuPont\u00ae 5142 gold paste for working electrode, where the electrochemical detection takes place; II) Modified DuPont\u00ae 6146 silver/palladium paste to obtain silver-chlorine-film reference electrode (Ag/AgCl), a stable and reference potential inside of the sensor\u2019s environment and finally III) ESL 5545-LS platinum paste for auxiliary or counter electrode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000102_s1052618808020040-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000102_s1052618808020040-Figure2-1.png", + "caption": "Fig. 2.", + "texts": [ + " An increase in the tilt angle and the vibrotransportation velocity is equivalent to a decrease in the effect of Coulomb friction on the vibrations of the pan and a decrease in friction coefficient f in (3), (4). If the vibration amplitude is relatively large, A* \u03c90V, cos[ /\u03c90A*)] \u2248 1, there is no such decrease and vibration amplitude A* is the same as that of the horizontal pan in (3), (4). In this case, the effect of vibrations on mass m is similar to linear viscous friction ( /\u03c90A*) \u2248 /\u03c90A*) in the third equation in (3). The amplitude\u2013frequency characteristic of pan (4) is symmetrical with respect to the axis \u03c9 = \u03c90 (Fig. 2, curve 1). Condition (5) corresponds to curve 2. The segments of A(\u03c9) that do not satisfy condition (5) with boundary points A and B are shown with a broken line. The stability of stationary regimes (4) is determined by means of equations for variations \u03b4 of system (3) (A = A* + \u03b4A, \u03d5 = \u03d5* + \u03b4\u03d5, and = V + ) whose coefficients are elements of square matrix ||bji ||. The x A \u03c8, x\u0307cos \u03c90A \u03c8;sin\u2013= = A\u0307 \u2013\u03b5 2\u03c90( ) 1\u2013 2\u03c90bA 4sf \u03c0 1\u2013 \u03b1cos( ) y\u0307/\u03c90A( )arcsin[ ]cos e \u03d5sin\u2013+{ },= \u03d5\u0307 \u03c9 \u03c90\u2013 \u03b5e 2\u03c90A( ) 1\u2013 \u03d5, y\u0307\u0307cos+ \u03b5g \u03b1sin 2 f \u03c0 1\u2013 \u03b1cos( ) y\u0307/\u03c90A( )arcsin\u2013[ ],= = \u03c90A y\u0307, A 0", + " GERTS value of each coefficient is determined from the partial derivative of the right-hand part of the jth equation in (3) by the ith variable in A, \u03d5, and at stationary points (4) (6) The necessary and sufficient conditions of asymptotic stability follow from the characteristic equation of the linearized system with elements (4) \u03bb3 + B1\u03bb2 + B2\u03bb + B3 = 0. In accordance with the Routh\u2013Hurwitz criterion, these conditions have the following form: (7) After transformations, conditions (7) taking into account (3) correspond to the inequality cos\u03b1 > 0, which is satisfied due to the existence of solutions of (4). Therefore, entire resonance curve 1 (Fig. 2) is stable, including the segments that do not satisfy inequality (5), which is not related to the conditions of existence of regimes (4). 3. In the case of zero initial conditions x = 0, = 0, and = 0 (A = 0 and \u03c8 = 0) and the effect of harmonic excitation E in the system (Fig. 1), both masses first move together. The parameters of the transient and stationary regimes are determined in this case from exact formulas for the linear oscillator with mass M + m and damping b. Using as an example (Fig. 2) two main cases of the amplitude\u2013frequency characteristics of this system, curves 3 show the parameters of the stationary regimes without sliding \u03c901 = K/(M + m). Curves 2 determine in this case the onset of sliding according to (5). We now consider a transient process denoted by thin arrows with points C and D (Fig. 2a). The vibration amplitude increases to point C, where the conditions for the onset of sliding may be determined exactly for system of equations (1). The process is further controlled by truncated equations (3) and ends at stable sta- y\u0307 b11 \u2013b 2sf \u03c0 1\u2013 \u03c90 3\u2013 YV 2 A* 3\u2013 , Y\u2013 1 V /\u03c90A*( )2 \u2013[ ] 0,5\u2013 \u03b1, b12cos e \u03d5*,cos= = = b13 = 2sf \u03c0 1\u2013 \u03c90 3\u2013 VYA* 2\u2013 , b21 = \u2013s 2\u03c90( ) 1\u2013 A* 2\u2013 \u03d5*, b22cos = \u2013e 2\u03c90( ) 1\u2013 A* 1\u2013 \u03d5*,sin b31 2gf \u03c0 1\u2013 \u03c90 1\u2013 VYA* 2\u2013 , b33 \u20132gf \u03c0 1\u2013 \u03c90 1\u2013 YA* 1\u2013 , b23 0, b32 0.= = = = B1 0, B3 0, B1B2 B3\u2013 0, B1> > > \u2013 b11 b22 b33+ +( ),= B2 b11b22 b11b33 b22b33 b23b32\u2013 b12b21\u2013 b13b31,\u2013+ += B3 b11b23b32 b12b21b33 b13b22b31 b11b22b33\u2013 b12b23b31\u2013 b13b21b32", + " Returning to the transient process, after point E, the vibration amplitude decreases according to (3) and tends to point F until condition (5) is violated. After this, the system loses one degree of freedom and transforms into a linear oscillator that tends again to the unrealizable stationary regime (the broken line on curve 2). This process is shown schematically in Fig. 3b, where A3 is the amplitude at point E. Similar regimes exist for frequencies larger than those at point B (the broken line on curve 1 in Fig. 2a). For description of these regimes, coordinates rather than their envelopes (3) must be known, and, therefore, it is not possible within the framework of this study. Nevertheless, it is proven that vibrotransportation is possible within broader ranges of amplitudes and frequencies than in the classical theory. The described regimes are characterized by \u201csoft\u201d excitation of vibrations, i.e., by a \u201csoft\u201d regime of vibrotransportation. The regimes with frequencies smaller than those at point G correspond to vibrations of an oscillator without vibrotransportation. The exact solution shows that, above and close to boundary (5), only regimes with finite-length stops in relative motion are possible. The boundary of existence of regimes with only instantaneous stops may be easily found from the formulas reported in [1] or from the diagram presented in [7] (8) where k = const is found using two parameters f and \u03b1 in [1]. According to (8), curve 4 (Fig. 2) always lies above curve 2 (shown by a dot-dash line). These stops are possible in segments AH and IB (Fig. 2a) of resonance curve 1 between curves 2 and 3. A \u201csoft\u201d regime is realized, which is similar to that in Fig. 3b, with vibrotransportation and vibration amplitudes located between resonance curve 1 and boundary 2 (thin arrow). If excitation amplitude E0 decreases, a situation is possible in the system shown in Fig. 2b where curve 2 passes above oscillator amplitudes (curve 3). The onset of regimes with vibrotransportation is possible only on the AB segment of curve 1 through \u201chard\u201d excitation of vibrations. The initial conditions for this regime are set by the corresponding selection of A(0), (0), \u03d5(0) \u2192 \u03c8(0) and inverse transform according to (2) to initial coordinates in (1). On segment HI of curve 1, stable vibrotransportation regimes are implemented, and on segments AH and BI, regimes of the type shown in Fig. 3b are implemented. \u201cIsolated\u201d segments AB of resonance curve 1 may also be reached by \u201cextending\u201d the vibration regime by increasing and subsequently slowly decreasing the excitation amplitude. Resonance curve 3 of the linear oscillator may be implemented in full as in Fig. 2b; therefore, it is rather difficult to find the \u201chard\u201d vibrotransportation regimes in an experiment without calculations. If the excitation amplitude is decreased further, curve 1 goes down (Fig. 3b) and has only one common point J with known amplitude and frequency \u03c90 on curve HI. Therefore, the possible regulation of vibrotransportation by changing the voltage amplitude E0 is limited according to the first expression in (4). 4. We now consider specific features of the obtained regimes. If mass m is fairly small (m/M ~ \u03b5), the term with Coulomb friction in the first equation in (1) will be of the second order of smallness and, in the first approximation, the mass motion does not affect vibrations", + " Therefore, the comparable values of the masses and the instantaneous change in the vibrating mass during the first impact play the key role in the appearance of the hard regimes of vibrotransportation. Taking into account the \u201cdecreasing\u201d characteristic of the drive results in an increase in the linear damping factor A* k\u03c9 2\u2013 ,> y\u0307 122 JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 37 No. 2 2008 M.E. GERTS, M.M. GERTS although it may also affect the mutual positions of resonance curves 1 and 3 in Fig. 2. The case of a relatively small transported mass is the only case where all of the results of the classical vibrotransportation theory are valid [4]; however, they hold in the first approximation alone. In the general case, the classical theory may be used to adjust the vibrotransportation velocity for nonstop regimes [1]: (9) In the second formula of (4), \u03c90 may be replaced without loss of accuracy by \u03c9 since in the third expression in (3) this replacement holds with an accuracy of \u03b52 owing to the formulas \u03c90 = \u03c9 \u2013 \u03b5\u2206 and /\u03c90A) = /\u03c9A) + \u03b5\u2206(\u2026) + \u03b52\u22062(\u2026) + \u2026", + " 3b is possible may not be recommended for practical use due to their possible stochasticity although it frequently appears in different systems between unstable limiting cycles rather than between stable ones. The regimes with hard excitation ensure the lowest voltage and amperage on the vibration exciter winding. They may be recommended for continuous vibrotransportation. The regimes with soft excitation are also suitable for dosing operations since they may be implemented when a mechanism is switched on. It should be noted that the range of vibrotransportation velocities is limited in both cases due to the limited character of the frequency ranges shown in Fig. 2, which is absent in the classical theory. ACKNOWLEDGMENTS This study was supported by the Russian Foundation for Basic Research, project no. 05-08-50183. REFERENCES 1. Blekhman, I.I. and Dzhanelidze, G.Yu., Vibratsionnoe peremeshchenie (Vibration Displacement), Moscow: Nauka, 1964. 2. Ul\u2019trazvuk. Malen\u2019kaya entsiklopediya (Ultrasound. Small Encyclopedia), Golyamina, I.P., Ed., Moscow: Sovetskaya Entsiklopediya, 1979. 3. Gerts, M.E. and Gerts, M.M., Synthesis of Self-Resonance Machines, Probl. Mashinostr" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002060_j.promfg.2018.06.045-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002060_j.promfg.2018.06.045-Figure4-1.png", + "caption": "Fig. 4: Concept of a reconfigurable four-pole permanent magnet spindle motor with switchable number of parallel branches", + "texts": [ + " The number of turns per phase is a significant parameter for the maximal achievable rotational speed, which is defined during the design process of the spindle motor. Usually, it cannot be changed after the machine has been manufactured. To enable the cost-efficient use of standard frequency inverters for HPC and HSC processes, the new spindle motor is designed with a reconfigurable winding system. A switchable number of parallel branches, exemplary shown for phase a of a star winding configuration in Fig. 4 (b), allows the change of the motor characteristics and enables a transition from HPC to HSC behavior by reducing the number of turns per phase. Assuming a constant line current, the maximum rotational speed is doubled and the maximum torque is halved, when the motor is switched to the next higher number of parallel branches. Requirements for a reconfigurable winding are power electronic or mechanical switches and a controlling logic, which prevents short circuits and initiates the switching process", + " Considering the available space inside a milling spindle and the torque requirements for typical HPC processes with up to \ud835\udc40\ud835\udc40shaft = 300 Nm, \ud835\udc5d\ud835\udc5d = 2 is chosen. Standard tools for HSC operations are solid carbide end mills. For the experimental investigations, a HSC-tool (Seco JS534160D1B.04ZNXT) with a diameter of 16 mm was used. The process parameters were specified according to tool manufacturer\u2019s specifications in order to avoid decreased tool lifetime. The cutting power measurements during cutting the aluminum alloys are presented in Fig. 4. The electrical power was measured with a ZES Zimmer LMG 500. The theoretical inner torque without losses is calculated based on the power equation (1). \ud835\udc40\ud835\udc40 = \ud835\udc43\ud835\udc43el 2 \u22c5 \ud835\udf0b\ud835\udf0b \u22c5 \ud835\udc5b\ud835\udc5b (1) With a terminal voltage of 360 V, the spindle exhibits a maximum power of 80 kW. This performance is sufficient according to the cutting power measurement results, shown in Fig. 5. Spindles with a switchable operation mode extending the working range of the rotational speed for HPC and HSC milling have a high potential to increase the economy of universal machine tools" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003000_ecc.2018.8550336-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003000_ecc.2018.8550336-Figure1-1.png", + "caption": "Fig. 1. Schematic adopted from [7] of the (a) spacecraft including three axisymmetric reaction wheels and four electric thrusters, and (b) Northfacing boom-thruster assembly. The first, second, and third axes of each reference frame are respectively denoted by red, green, and blue vectors.", + "texts": [ + " The anti-symmetric projection operator Pa(\u00b7) : R 3\u00d73 \u2192 so(3), is given by Pa(U) = 1 2 ( U \u2212 UT ) , for all U \u2208 R 3\u00d73, where so(3) = {S \u2208 R 3\u00d73 | S + ST = 0}. The cross operator, (\u00b7)\u00d7 : R 3 \u2192 so(3), is defined as a\u00d7 = \u2212a\u00d7 T = \u23a1 \u23a3 0 \u2212a3 a2 a3 0 \u2212a1 \u2212a2 a1 0 \u23a4 \u23a6, where aT = [ a1 a2 a3 ] . The uncross operator, (\u00b7)v : so(3) \u2192 R 3, is defined as Av = [ a1 a2 a3 ]T , where A = a\u00d7. The physical vector describing the position of p relative to q is r\u2212\u2192pq . Similarly, the angular velocity of Fb relative to Fa is \u03c9\u2212\u2192ba. Consider the satellite shown in Fig. 1, which consists of a rigid bus equipped with three axisymmetric reaction wheels and four electric thrusters mounted on gimbaled booms, which is nominally in a circular GEO orbit. The control objectives are to 1) minimize the effect of quantization on \u0394v and 2) limit the number of on-off thruster pulses, while ensuring that the satellite is maintained within the prescribed station-keeping window, a nadir-pointing attitude is maintained, angular momentum stored in the reaction wheels is unloaded, and the limitations of the thrusters (e", + " An MPC policy is also adopted in this paper, with a novel quantization scheme that is designed to specifically minimize the effect of quantization on \u0394v and reduce the number of on-off thruster pulses compared to [7]. The Earth-centered inertial (ECI) frame is defined as Fg . The reference frame Fp is aligned with the spacecraft bus, where nominally p\u2212\u2192 1 points towards the Earth and p\u2212\u2192 2 points North. The angular velocity of Fp relative to Fg is \u03c9\u2212\u2192pg and the DCM describing the attitude of the spacecraft (i.e., Fp) relative to Fg is Cpg . The spacecraft center of mass is denoted by point c in Fig. 1(a). The position of c relative to a point w at the center of the Earth is given by r\u2212\u2192cw. The equations of motion of the satellite are [7] r\u0308cwg = \u2212\u03bc rcwg\u2225\u2225rcwg \u2225\u22253 + apg + 1 mB CT pgfthrust p , (1a) JBc p \u03c9\u0307pg p = \u2212\u03c9pg\u00d7 p ( JBc p \u03c9pg p + Js\u03b3\u0307 ) \u2212 Js\u03b7 + \u03c4p p + \u03c4 thrust p , (1b) C\u0307pg = \u2212\u03c9pg\u00d7 p Cpg, (1c) \u03b3\u0308 = \u03b7, (1d) where mB is the mass of the spacecraft, JBc p is the moment of inertia of the spacecraft relative to point c and resolved in Fp, \u03b3T = [ \u03b31 \u03b32 \u03b33 ] are the reaction wheel angles, \u03b7 is the acceleration of the reaction wheels, Js is the moment of inertia of the reaction wheel array, fthrust p is the force produced by the thrusters, \u03c4 thrust p is the torque produced by the thrusters, apg includes acceleration perturbations, and \u03c4 p p includes torque perturbations. The thruster configuration is illustrated in Fig. 1, where four electric thrusters are mounted on two boom-thruster assemblies, which nominally point North and South, respectively. A detailed view of the North- facing boom-thruster assembly is in Fig. 1(b). Each assembly has two fixed gimbal angles, \u03b1\u0304a and \u03b2\u0304a, a \u2208 {n, s}, as well as an actuated gimbal angle \u03b3a, a \u2208 {n, s}. The subscripts n and s refer to the North- and South-facing assemblies, respectively. The position of the actuated gimbal of thruster i relative to the spacecraft center of mass is r\u2212\u2192qic. The thrusters are canted by fixed angles \u03b4i, i = 1, 2, 3, 4, such that for \u03b3\u0304a, a \u2208 {n, s}, each thruster fires through the spacecraft center of mass. The force vector produced by thruster i is f\u2212\u2192 i, and is resolved in Fp as fip = \u2212f iCT ipC2(\u03b3a)13, where f i = \u2223\u2223\u2223 f\u2212\u2192i \u2223\u2223\u2223 is the thrust magnitude, 13 = [ 0 0 1 ]T , Cip = CiaCap, Cia = C1(\u03b4i)C2(\u03b2\u0304i)C3(\u03b1\u0304a), Cnp = C3(\u03c0), and Csp = C1(\u03c0)C3(\u03c0)", + " The control input sequence is u(t + j) = u\u2217 j|t, j = 0, . . . , Nfb \u2212 1, where U\u2217 t is the minimizer of (2), and Nfb is the number of time steps between control updates. IV. SINGLE-PULSE QUANTIZATION SCHEME The low-thrust electric thrusters considered for this space- craft are operated with on-off pulses. The control input generated by the MPC policy described in Section III-D is a continuous thrust value for each thruster, which cannot be used directly with on-off thrusters, or in the propulsion system assembly shown in Fig. 1. As such, the control input must be quantized to on-off pulses that satisfy the physical constraints of the thrusters and the propulsion system assem- bly. A PWM quantization scheme is developed in [7] with a fixed frequency of five on-off pulses per time step with varying pulse widths such that the average thrust matched the constant thrust of the MPC control input over each time step. This works well, but leads to a large number of on-off pulses, on the order of 30 pulses per thruster per orbit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001574_1.5018474-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001574_1.5018474-Figure2-1.png", + "caption": "FIG. 2. Schematic of the experimental facility.", + "texts": [ + " The present experiments were conducted in a horizontal shock tube with an inner diameter of 200 mm, as shown in Fig. 1. The lengths of its driver and driven sections are 1260 and 2000 mm, respectively. A test section made of a transparent acrylic material has a 200 mm 200 mm square cross-section and a 2000 mm axial length. A round-to-square transitional section with a length of 300 mm was adopted to connect the test section and the driven section. However, the outlet of the flow channel of the test section is directly connected with the atmosphere. Figure 2 shows the schematic of the experimental facility. In the present experiments, aluminum diaphragms with thicknesses varying from 0.05 to 0.15 mm were employed to generate planar shock waves. A dynamic pressure transducer connected with a digital displayer for pressure was located at a position of 720 mm from the end-wall of the driver section to monitor the inside pressure. Other two piezoelectric dynamic pressure transducers (CY-YD-205, measurement range of 0\u201330 MPa) denoted by T1 and T2 in Fig. 2 were mounted on the bottom wall of the test section to measure the velocities of incident shock waves. They were connected to a high-speed data acquisition system (TST5510, maximum sampling frequency of 10 MHz). The data acquisition in the present tests was triggered with a pressure signal generated by the arrival of an incident shock wave at transducer 1. Figure 3 shows two typical step-like pressure wave patterns measured by transducers 1 and 2. The variables, L12 and Dt, denote the distance between the two transducers and the interval time-of-arrival of the incident shock wave at the two transducers, respectively", + " Besides, large damped oscillations can be observed in each pressure signal following the timeof-arrival of the incident shock wave. It is believed that these oscillations are caused by the pressure transducer not being perfectly flushed with the inside wall of the test section. The imaging of moving shock and compression waves in this study was achieved by using a high speed Schlieren system. This system consists of a set of split-type Schlieren instrument, a high speed camera (FASTCAM SA5, maximum shooting speed of 1 106 fps), and a dedicated computer, as shown in Fig. 2. The Schlieren instrument includes a light source system, two 400 mm diameter concave reflectors 1 and 2 (focal length of 4000 6 10 mm, accuracy of 1/20 wavelength of light) and an imaging system. Here, the light source is a halogen lamp with a variable power of 0\u201320 W. In practice, a power of around 10 W was set to obtain a relatively good image quality. Besides that, a plane mirror with a diameter of 400 mm was used to solve the problem of space limitation of the laboratory. The parameter settings including a shooting speed of 100 000 fps and an exposure time of three millionths of a second were adopted for the high-speed camera in the experiments", + " 33 was used in this study. As shown in Fig. 4(a), each of the spheres with a diameter of D\u00bc 40 6 0.1 mm is assembled from two hollow aluminumalloy hemispheres with screw threads. The special design of the internal structure of the sphere model permits one piezoelectric accelerometer (TST266A01, measurement frequency range of 1\u2013100 kHz, degree of nonlinearity of less than 0.5%) being perfectly mounted inside each sphere. All the accelerometers are connected to the high-speed data acquisition system (shown in Fig. 2). In order to control the distance between spheres and construct a multi-sphere model, for example, a double-sphere model inside the test section [shown in Fig. 4(b)], four rows of small through-holes with a diameter of 1 mm are placed on each sidewall of the test section. Meanwhile, one column of through-holes is arranged on each of the top and bottom walls along the respective central axes. Each sphere has a 1 mm diameter through-hole that allows the sphere to be fastened at an expected position inside the test section using a 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003710_978-3-030-13317-7_8-Figure8.4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003710_978-3-030-13317-7_8-Figure8.4-1.png", + "caption": "Fig. 8.4 Set of worn gears considered for this proposed analysis. a Healthy and b 25%, c 50% and d 75% of uniform wear", + "texts": [ + " Indeed, the analyzed gearbox wear consists of a uniform wear which is similar to that produced in gearboxes after a long useful life in real industrial applications. Accordingly, three different levels of wear in gears are evaluated to analyze the nonlinear effects that a typical fault in gears produce in the vibration response and in the stator current consumption. The 4:1 ratio gearbox consists of two gears, the driver gear has 18 teeth and the driven gear has 72 teeth; thus, the wear has been artificially induced uniformly by a gear factory in all teeth of three similar driven gears. From Fig. 8.4a\u2013d are shown the set worn gears analyzed in this proposal: healthy and 25, 50 and 75% of uniform wear, respectively. In regard with the damaged bearing, the IM uses a bearing model 6205-2ZNR that is located in rotor on the output shaft side; thus, a similar bearing element has been also artificially damaged by means of drilling a through-hole on its outer race with a tungsten drill bit of 1.191 mm diameter to produce the bearing with damage. In Fig. 8.5 is shown a picture of the damaged bearing used to perform the proposed analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002890_j.matpr.2018.08.148-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002890_j.matpr.2018.08.148-Figure1-1.png", + "caption": "Fig. 1. General view of a leaf spring.", + "texts": [ + "/ Materials Today: Proceedings 5 (2018) 26760\u201326765 26761 Decisive influence on the type of the system has a vehicle\u2019s purpose. In the military vehicles, suspension systems based on the leaf springs prevail significantly. A research object, subjected to the numerical and experimental analysis, was the semi-elliptical spring. This element is utilized in heavy-duty vehicles, operating in a very difficult site conditions. It is used as a spring element of the suspension, mounted on the frontal axis of three and four axle vehicles. A general view of the analysed object is shown in figure 1. It is comprised of the nine leaves with a variable length. The value of a nominal stiffness of the examined spring was defined and it was around 330N/mm. According to Polish standard [5] considering testing the springs of the suspension system, the semi-elliptical spring should withstand more than 200 000 cycles with constant deflection around 100mm. The spring manufacturers are obligated to perform this type of tests before introduction the product to the market, so the springs are designed to meet the regulations requirements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002206_s1068798x18070043-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002206_s1068798x18070043-Figure11-1.png", + "caption": "Fig. 11. Cutting process in the zone of cutter exit from the blank.", + "texts": [ + " Since crack formation along the conditional shear plane is typical of cutting, the chip will separate from the front surface when the cutter is at some distance from the end of the black, on account of the crack that forms (Fig. 10e). As a result, the forces between the cutter and the blank will be due to contact of the machined surface with the tool\u2019s rear surface, and cutting will stop (Fig. 10f). We now consider the cutting process at the edge in the moment of chip breakaway from the blank and the cutter\u2019s front surface. The cutting scheme for this moment is shown in detail in Fig. 11, which is valid for both positive and negative values of the rake angle \u03b3. To simplify the analysis, we assume a plane strain state. In other words, we assume zero strain perpendicular to the plane of the drawing. In addition, we make the following assumptions. The cutter is introduced in the blank at speed v0. The cutter\u2019s velocity vector v0 is assumed perpendicular to the end of the blank (segment AB), unless that proves inconsistent with the conditions of the calculation. The zone of intense plastic deformation that appears is enclosed within the volume bounded by the cutter surface (segment OC), the surface of the blank\u2019s free end (segment AB), and segments AO and BC", + " When the zone of plastic1 For parts 1 and 2, see Steel Transl., nos. 5 and 6, 2018. 519 deformation (height h4) is to the right of the cutter tip (point O), elastic unloading occurs, accompanied by contact with the cutter\u2019s rear surface over length l3 [6]. Consider the boundary conditions for the tangential stress, which are analogous to those in Vorontsov\u2019s work but written in more general form where yf is the y coordinate at the edge of the plasticdeformation zone (segment AB). It follows from Fig. 11 that yf = f(x). \u03c4 = \u03bc\u03b2\u03c3 =\u23a7 \u23a8\u03c4 = \u2212 \u03b2\u03c3 =\u23a9 f , when 0; 0.5 , when ; xy s xy s y y y RUSSIAN The stress along the x axis may be obtained in the following form after the mathematical procedures outlined in [6] (7) Compressive stress due to the contact-friction force over length l acts along the x axis from zone BCDK to the upper boundary of the zone of intense plastic deformation (segment BC). The mean (over the chip thickness) frictional stress is The contact length of the chip with the tool\u2019s front surface is [7] (8) The term uh1(1/(ucos \u03b3 \u2013 1) is negligibly small" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002595_chicc.2018.8482808-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002595_chicc.2018.8482808-Figure3-1.png", + "caption": "Fig. 3: Geometric relationship of the semi-strapdown seeker\u2019s angles The control objective of the coordinator stable tracking loop is to make dOX axis follow sOX axis and to make tracking error to be arbitrarily small. According to the geometric relation, it is obtained", + "texts": [ + " The dynamics model of the missile can be described as follows: cos 1 cos m mz m m z mz z mz m m m m g L V mV M J L mg mV (3) where is the angle of attack, mz is the pitch angle rate, zM is the pitching moment, zJ is the moment of inertia about the z axis, is the pitch angle. Based on the theory of missile aerodynamics, the lift force L and the pitching moment zM can be modeled as 2 z mz z L L z z z z mz z z M L QsC QsC Qsl mM Qslm Qslm V (4) where z is the deflection angle of the control surface, 2 / 2mQ V is the dynamic pressure, s is the reference area, l is the reference length, LC and z LC are two lift coefficient, zm , mz zm and z zm are three pitching moment coefficient. 2.3 Seeker angle tracking model The spatial angle relation of the seeker is shown in Fig. 3, where is the gimbals angle, is the angle between the center line of the detector and the base line, is the misalignment angle. =q (5) Then one gets = q mz (6) where is the seeker gimbals angle rate. To facilitate the analysis and design, the longitudinal guidance and control system of the missile is simplified, and the following assumptions are made. Assumption 1: The lift force produced by the control surface is much smaller than the lift force, for the little control surface, that is 0z LC ; Assumption 2: At the terminal guidance stage, the velocity of the missile is constant, that is, mV is a constant value; Assumption 3: At the terminal guidance phase, the engine does not work, that is, 0P " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002348_978-981-13-0212-1_34-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002348_978-981-13-0212-1_34-Figure7-1.png", + "caption": "Fig. 7 The type-2 defect bearing", + "texts": [], + "surrounding_texts": [ + "To implement the proposed technique of fault diagnosis an experimental setup is developed. It consists of a single phase induction motor. The experimental setup and its model are shown in Figs. 3 and 4 respectively. To carry out the experiment three bearings are taken; one healthy bearing and two different defective bearings. In the type-1 defect bearing there is a cut at the outer race and in the type-2 defect bearing there is a hole at the outer race of the bearing. The healthy, type-1 and type-2 defect bearings are shown in Figs. 5, 6, and 7 respectively. The setup also Fig. 4 The experimental setup model Fig. 5 The healthy bearing Fig. 6 The type-1 defect bearing consists of a data acquisition system to acquire the vibration signal from the bearings. The data acquisition system consists of an accelerometer (PCB 325c-03), which is a vibration sensor, a 4-channel DAQ card (NI-9234) and a PC with LabVIEW software. The specification for the motor and for the bearings are shown in Tables 2 and 3 respectively. Initially the healthy bearing is mounted on the shaft of the motor and the corresponding vibration signature is acquired. Then the type-1 and type-2 defect bearings are mounted and the corresponding vibration signals are acquired. ANC using an EMD algorithm is implemented on all the acquired vibration signals and the filtered vibration signals from the healthy, type-1 and type-2 defect bearings are shown in Figs. 8, 9, and 10 respectively. Table 2 Bearing specification Model 6203z No of balls (N) 8 Ball diameter (D) 6.74 mm Pitch diameter (D) 30 mm Contact angle (\u03b1) 0 Table 3 Motor specification Type Induction motor RPM 8 Power 0.5 hp AMP 2 AMP Make Crompton" + ] + }, + { + "image_filename": "designv11_92_0000964_med.2009.5164601-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000964_med.2009.5164601-Figure1-1.png", + "caption": "Fig. 1. Pictorial Representation of inner and outer parameter hopping", + "texts": [ + " The modified weight updating laws when using parameter \u201chopping\u201d are given as follows: a) For the updating of the weights W i l W\u0307 i l = \u2212\u03b31 ( Xi )T \u03beisl(x) if Xi \u00b7 W i l \u2208 P2 or Xi \u00b7 W i l = \u00b1\u03b5i andXi \u00b7 W\u0307 i l >< 0 \u2212\u03b31 ( Xi )T \u03beisl(x)\u2212 2 ( XiW i l (Xi) T ) tr{(Xi)T Xi} otherwise (27) where P2 = {W i i,l : |Xi \u00b7 W i l | \u2264 \u03b5i}. b) For the updating of the weights 1W i 1W\u0307 i = \u2212\u03b32 ( 1Xi )T \u03beiuisi(x) if 1Hi \u2208 P1 or 1Hi = \u00b1\u03b8i or 1Hi = \u00b1\u03c1i and 1H\u0307i <> or >< 0 \u2212\u03b32 ( 1Xi )T \u03beiuisi(x)\u2212 2 ( 1Xi1W i(1Xi) T ) tr{(1Xi)T 1Xi} otherwise (28) where 1Hi = 1Xi \u00b7 1W i and P1 = {1W i : \u2223 \u2223 1Xi \u00b7 1W i \u2223 \u2223 \u2264 \u03c1i and \u2223 \u2223 1Xi \u00b7 1W i \u2223 \u2223 \u2265 \u03b8i}. This procedure is depicted in Fig. 1, in a simplified 2- dimensional representation. Lemma 1: Based on the adaptive laws (27), (28) the additional terms introduced in the expression for V\u0307 , can only make V\u0307 more negative. Proof: Let that 1W \u2217i contains the actual unknown values of 1W i such that \u2223 \u2223 1Xi \u00b7 1W \u2217i \u2223 \u2223 >> \u03b8i and that 1W\u0303 i = 1W i \u2212 1W \u2217i. Then, the weight hopping can be equivalently written with respect to 1W\u0303 i as \u22122\u03b8i 1W\u0303 i/\u20161W\u0303 i\u2016. Under this consideration the modified updating law is rewritten as 1W\u0307 i = \u2212\u03b32 ( 1Xi )T \u03beiuisi1(x) \u2212 2\u03b8i 1W\u0303 i/\u20161W\u0303 i\u2016" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001009_iciea.2008.4582880-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001009_iciea.2008.4582880-Figure2-1.png", + "caption": "Fig. 2. Doubly-fed generator stator flux oriented vectors frame", + "texts": [ + " After 3s/2r coordinate transformation, the basic equations of the generator models in two-phase rotate d-q coordinate are shows as follows\uff1a voltage equations: sd s sd sd s sq sq s sq sq s sd rd r rd rd sl rq rq r rq rq sl rq u R i p u R i p u R i p u R i p \u03c8 \u03c9 \u03c8 \u03c8 \u03c9 \u03c8 \u03c8 \u03c9 \u03c8 \u03c8 \u03c9 \u03c8 = + \u2212\u23a7 \u23aa = + +\u23aa \u23a8 = + \u2212\u23aa \u23aa = + +\u23a9 (3) flux equations: sd s sd m rd sq s sq m rq rd m sd r rd rq m sq r rq L i L i L i L i L i L i L i L i \u03c8 \u03c8 \u03c8 \u03c8 = +\u23a7 \u23aa = +\u23aa \u23a8 = +\u23aa \u23aa = +\u23a9 (4) torque equation: ( )e p m sq rd sd rqT n L i i i i= \u2212 (5) In these equations, ,sd squ u are respectively stator\u2019s voltage d,q-axis component \uff0c ,rd rqu u are respectively rotor\u2019s voltage d,q-axis component, ,sd sqi i are respectively stator\u2019s current d,q-axis component \uff0c ,rd rqi i are respectively rotor\u2019s current d,q-axis component, ,sd sq\u03c8 \u03c8 are respectively stator\u2019s flux d,q-axis component \uff0c ,rd rq\u03c8 \u03c8 are respectively rotor\u2019s flux d,q-axis component\uff0c sR , sL are respectively stator wings\u2019 resistance and self-inductance \uff0c rR , rL are respectively rotor wings\u2019 resistance and self-inductance, mL is the mutual- inductance between stator and rotor, s\u03c9 is the synchronous angular speed, sl s r\u03c9 \u03c9 \u03c9= \u2212 is slip angular speed[3]. When we put the d-axis of the d-q coordinate system on the doubly-fed generator\u2019s stator flux direction(as in figure 2 shows), the q-axis component of the stator flux is 0. So we can get follows[4]: 0 sd s sq \u03c8 \u03c8 \u03c8 =\u23a7\u23aa \u23a8 =\u23aa\u23a9 (6) We can get follows from the model equations: 0 sd s sd m rd s m ms sq s sq m rq L i L i L i L i L i \u03c8 \u03c8 \u03c8 = + = =\u23a7\u23aa \u23a8 = + =\u23aa\u23a9 (7) In this equation set, msi is the generalized excited current; So we can get follows: ( )m sd ms rd s m sq rq s Li i i L Li i L \u23a7 = \u2212\u23aa\u23aa \u23a8 \u23aa = \u2212 \u23aa\u23a9 (8) Pg 2046 Then the rotor flux equations change to follows: 2 2 2 ( ) ( ) m m rd ms r rd s s m rq r rq s L Li L i L L LL i L \u03c8 \u03c8 \u23a7 = + \u2212\u23aa \u23aa \u23a8 \u23aa = \u2212\u23aa\u23a9 (9) From these equations we can obtain the torque equation as follow: 2 m e p ms rq s LT n i i L = \u2212 (10) Because the voltage drop on the stator resistance is far less than the power grids voltage, it can be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000752_sice.2008.4655087-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000752_sice.2008.4655087-Figure3-1.png", + "caption": "Fig. 3 Coordinate system in the single leg stance period.", + "texts": [ + " Output torque at each joint is calculated using simple PD control expressed by the following equation. This PD control is used excluding the knee joint of the swing leg in the stance-swing state. \u03c4 ph j = \u2212k ph j p(\u03b8 ph j c \u2212 \u03b8 ph j d) \u2212 k ph j v(\u03b8\u0307 ph j c \u2212 \u03b8\u0307 ph j d) (1) \u03c4 ph j , \u03b8 ph j c and \u03b8 ph j d are output torque, measured joint angle and desired joint angle at a joint j \u2208 {h, k, a} in PR0001/08/0000-2507 \u00a5400 \u00a9 2008 SICE - 2508 - a phase ph \u2208 {sw, st}, respectively, where h, k and a mean hip, knee and angle joint, and sw and st mean swing and stance phase, respectively (Fig. 3). \u03b8\u0307 ph j c and \u03b8\u0307 ph j d are measured and desired joint angular velocity. In this study, \u03b8\u0307 ph j d was set to zero. The treadmill has two belts equipped with a DC motor for each, so that the speed of each belt (i.e., each leg) can be controlled independently. In different test stages, Tetsuro walks on the treadmill with the two belts either moving at the same speed (\u201ctied\u201d configuration) or different speed (\u201csplitbelt\u201d configuration). We can change the belt speed from 0 m/s to 0.6 m/s in every 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002153_978-3-319-96181-1_12-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002153_978-3-319-96181-1_12-Figure1-1.png", + "caption": "Fig. 1. Dynamic model of spur gearbox with an elastic coupling", + "texts": [ + " The system is powered by diesel engine and the second model of Nelson and Crandall (1992) is adopted for the elastic coupling. Excitations due to the fluctuation of load and speed of acyclism regime and eccentricity defect of gearbox are introduced to the dynamic model. The dynamic response is computed through Newmark algorithm and results are shown using Wigner\u2013Ville distributions. The studied system is composed by a diesel engine motor and a receiver which are connected through one stage of spur gearbox and an elastic coupling located between the motor and the pinion. Figure 1 show the corresponding dynamic model which is divided into three blocks. This model was proposed by Hmida et al. (2016, 2017). The pinion and the wheel of gearbox and the diesel engine motor are assumed as rigid bodies. Transmission shafts are assumed massless and have torsional stiffness Khi and torsional damping Chi (i = 1, 2, 3). They are supported by bearings which are modeled with parallel springs (Kxi, Kyi) and damping (Cxi, Cyi). The model of Nelson and Crandall (1992) is adopted for the elastic coupling because this model is best approach to describe the dynamics of elastic couplings (Tadeo and Cavalca 2003; Tadeo et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001062_cimca.2008.135-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001062_cimca.2008.135-Figure2-1.png", + "caption": "Figure 2. Platform and rod frames from a view angle with azimuth 130o and elevation 30o.", + "texts": [], + "surrounding_texts": [ + "benchmark problem from the plane to the three dimensional real world [2], [3]. The kinematics and dynamics of the nonholonomic wheeled mobile robots have been introduced and developed by many researchers [4], [5], [6]. However, as stated in [1], none of them can be directly applied to a mobile inverted pendulum stabilization problem with the constrained steering angle of a nonholonomic platform. The system stabilized in [1] has no platform, and the free movement of the pendulum is only in forward-backward direction due to the system structure. This paper contributes the kinematics and dynamics of a nonholonomic mobile platform for the dynamic stabilization of a freely moving inverted pendulum in a horizontal planar path. The dynamic modeling of the system is accomplished using homogeneous coordinate transformations with the Lagrangian method. An intermediary joint and link are required to represent the two independent rotation modes of the rod [12], but the Denavit-Hartenberg convention [11] is not applied to keep the z-axis of the of the platform, intermediary and rod frames directed upward. This paper is organized as follows: Section 1 describes the four-wheeled mobile inverted pendulum system. Section 2 presents the main components and kinematics relations. The dynamic equations of motion are developed in Section 3. Section 4 presents the conclusion and future research directions on FWMIP systems and the mobile inverted pendulum. 2. Kinematics of the System We represented the two rear wheels by a single central fixed wheel, and the two steering wheels by a single virtual central steering wheel similar to [9], such that the system is simplified to a bicycle in x-y plane [4]. Virtual central steering angle \u03d5 provides a central steering radius r = lp cot(\u03d5). The right and the left steering wheel angles \u03d5r = atan(lp /(r+wp/2)), and \u03d5l = atan(lp /(r\u2013wp/2)) are set by a steering mechanism depending on the front wheel distance wp.\n978-0-7695-3514-2/08 $25.00 \u00a9 2008 IEEE DOI 10.1109/CIMCA.2008.135\n409", + "We assigned coordinate frames Fp and Fr, to the platform and to the inverted pendulum rod to accomplish the kinematics and dynamics analysis of the system as seen in Figures 1 and 2. The x-axis of the platform frame is oriented toward the front of the platform, and z-axis is selected upwards. Accordingly, the nonholonomic constraint restricts the rear wheel motion in local y direction. The frame is placed at the joint point of the rod to achieve easy stabilization of transversal rod motion, although it is known that placing the platform frame at the central fixed wheel simplifies the kinematic relation regarding this frame [4]. The position of the platform is described by the vector\nqp= (x, y, \u03c8), where x, y are the coordinates of the platform frame in the base frame, and \u03c8 is the orienta-\nMovement with a steering angle \u03d5 =0 gives a linear motion along the x-axis of the platform, while \u03d5 0 results in a tangential motion \u0394d of the platform on a circle with the steering radius r = lpC\u03d5 /S\u03d5. This circular path turns the platform \u0394\u03c8 = \u0394d / r amount about the z-axis with respect to the initial platform position. Rotations of a coordinate frames about x, y and z-axes and its translation to a point p = (x, y, z)T can be expressed by homogeneous transformation matrices\nrot(x,\u03b1) = 1 0 0 0 0 C S 0 0 S C 0 0 0 0 1 \u03b1 \u03b1 \u03b1 \u03b1 \u2212 ; rot(y,\u03b2)= C 0 S 0 0 1 0 0 S 0 C 0 0 0 0 1 \u03b2 \u03b2 \u03b2 \u03b2\u2212 ;\nrot(z,\u03c8)= C S 0 0 S C 0 0 0 0 1 0 0 0 0 1 \u03c8 \u03c8 \u03c8 \u03c8 \u2212 ; trans(x,y,z)= 1 0 0 0 1 0 0 0 1 0 0 0 1 x y z ; (1)\nwhere \u03b1, \u03b2, \u03c8 are angular displacements, and the abbreviations C\u03b8 and S\u03b8 denote cos(\u03b8 ) and sin(\u03b8 ) [11]. The nonholonomic motion of the platform is conveniently expressed by a chain of homogeneous translation and rotation transformation matrices. The final platform location shown in Figures 1 and 2 is obtained relative to the initial platform frame by the following chain of transformation matrices 0Tp1 trans(\u2013 lp , lp cot(\u03d5) , 0 ) . rot(z, \u0394\u03c8) . trans(lp , \u2013 lp cot(\u03d5) , 0) ; (2) Fp1 p1Tp2 Fp2 (3)\nFrom Fp2 to Fp1 the transformation p1Tp2 is p1Tp2= ( )\n( ) C S 0 S C C S S / S S C 0 C C S S C / S\n0 0 1 0 0 0 0 1\np\np\nl\nl \u03c8 \u03c8 \u03d5 \u03c8 \u03d5 \u03c8 \u03d5 \u03d5 \u03c8 \u03c8 \u03d5 \u03c8 \u03d5 \u03c8 \u03d5 \u03d5 \u0394 \u0394 \u0394 \u0394 \u0394 \u0394 \u0394 \u0394 \u2212 + \u2212 \u2212 \u2212 . (4)\nFurther, the initial platform frame is transferred to the base frame Fp0 by rotation and translation transformations. 0Tp1 = trans(lp , 0 , 0) . rot(z,\u03c8) (5) F0 = 0Tp1 Fp1 . The overall transformation 0Tp is obtained as 0Tp = 0Tp1 p1Tp2 .\n= C S 0 (C C +(S S )C /S ) S C 0 (S S (C C )C /S )\n0 0 1 0 0 0 0 1\np\np\nl x l y \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03d5 \u03d5 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03d5 \u03d5 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 +\u0394 \u2212 \u2212 \u2212 + \u2212 \u2212 \u2212 + .\n(6)", + "Thus, the movement (\u0394x, \u0394y, \u0394\u03c8) between the final and initial positions of the platform for the steering wheel angle \u03d5 and the tangential path distance \u0394d is \u0394\u03c8 = (\u0394d / lp ) tan(\u03d5 ) \u0394x = lp (C\u03c8+\u0394\u03c8 \u2013 C\u03c8 + (S\u03c8+\u0394\u03c8 \u2013 S\u03c8) / tan(\u03d5) ) , \u0394y = lp (S\u03c8+\u0394\u03c8 \u2013 S\u03c8 + (C\u03c8+\u0394\u03c8 \u2013 C\u03c8) / tan(\u03d5) ) . (7)\nThe steering angle \u03d5 =0 gives divide by zero in (7) due to approach of r to infinity. In that case (4) converges to trans(\u0394d, 0, 0). Consequently, the overall transformation comes out\n0Tp= C S 0 C S C 0 S 0 0 1 0 0 0 0 1 d x d x \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u2212 \u0394 + \u0394 + , (8)\nresulting in incremental kinematics relations \u0394\u03c8 = 0; \u0394x = \u0394d C\u03c8 ; \u0394y = \u0394d S\u03c8 . (9)\nThe rod frame Fr is placed to the free end of the inverted pendulum rod, so that the frame position conveniently indicates deviation from the vertical pose. At the platform-rod joint the rod rotates freely about y and x axes. For a rigorous dynamic representation of this two degree-of-freedom joint we use an intermediary link and joint that provides rotation about x-axis. pTri = rot(x,\u03b1) ; (10) riTr = rot(y, \u03b2) trans(0, 0, lr) (11) In this configuration, eliminating the virtual joint and link restricts the motion of the rod only into x-z plane. The transformation\npTr = pTri riTr (12)\nFp = pTr Fr (13) converts the coordinates from the rod frame Fr to the platform frame Fp. The transformation pTrv and riTr are\npTri = 1 0 0 0 0 C S 0 0 S C 0 0 0 0 0 \u03b1 \u03b1 \u03b1 \u03b1\u2212 ; riTr = C 0 S S 0 1 0 0 0 C C 0 0 0 1 r r l S l \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 \u2212 \u2212 ,\nand the combined transformation pTr is\npTr = C 0 S S S S C S C S C C S S C C C C\n0 0 0 0\nr\nr\nr\nl l l \u03b2 \u03b2 \u03b2 \u03b1 \u03b2 \u03b1 \u03b1 \u03b2 \u03b1 \u03b2 \u03b1 \u03b2 \u03b1 \u03b1 \u03b2 \u03b1 \u03b2 \u2212 \u2212 \u2212 \u2212 . (14)\nThe coordinate transformation matrices 0Tp , pTri and riTr represent the position and orientation of the platform and rod to derive the dynamics of the system.\n3. Dynamics of the System Let q =(q1, \u2026 qn)T be vector of generalized joint displacements of an n degree-of-freedom open chain mechanism. The derivatives of Lagrangian L= K P give the generalized joint-force \u03c4i at the joint-i.\n\u03c4i = dt d iq \u2202 \u2202 L \u2212 iq \u2202 \u2202 L , (15)\nwhere K and P are the overall kinetic energy and potential energy of the analyzed frictionless system [11]. Integrating kinetic energy of differential masses over the complete mass of the ith link gives the kinetic energy ki of the link. ki =\ni im\nd k\n= 12Trace[(\u03a3i p=1 \u03a3i r=1 0 i pq \u2202 \u2202 T Jpi( 0 i rq \u2202 \u2202 T )T qr qp )], (16) where Jpi = mi\nirdm (irdm)T dmi is the pseudo-inertia matrix of the ith link. The potential energy pi of the ith link is pi = \u2212 mi ga T 0Ti icmi , (17) where ga is the gravitational acceleration vector, mi is the mass, and 0Ti\nicmi = 0cmi is the center of mass of the ith link. The time and joint variable derivatives of the Lagrangian give the generalized jointforce\n\u03c4i = dt d ( ) iq \u2202 \u2212 \u2202 K P \u2212 ( ) iq \u2202 \u2212 \u2202 K P , (18)\n= \u03a3n r=i \u03a3n k=1 ( 0 r kq \u2202 \u2202 T Jpr( 0 r iq \u2202 \u2202 T )T) q k\n+ \u03a3n j=1\u03a3n k=1\u03a3n r=1Trace( 2 0 r k jq q \u2202 \u2202 \u2202 T Jpr( 0 r iq \u2202 \u2202 T )T)q k q j\n\u2212 \u03a3n r=i mr ga T 0 r iq \u2202 \u2202 T rcmr . (19)\nFurther, (19) can be put in Uicker-Kahn form [11].\n\u03c4i =\u03a3n j=1 Dij qj +\u03a3n j=1 \u03a3n k=1 Cijk q j q k + g i . (20)\nwhere\nDij =\u03a3 n r=max(i,j) Trace( 0 r jq \u2202 \u2202 T Jpr ( 0 r iq \u2202 \u2202 T )T) ;\nCijk =\u03a3n r=max(i,j,k)Trace(\n2 0 r j kq q \u2202 \u2202 \u2202 T Jpr ( 0 r iq \u2202 \u2202 T )T), and\ngi = \u2212 \u03a3n r=i mr ga T 0 r iq \u2202 \u2202 T rcmr , (21)\nare the inertial, coriolis, and gravitational terms of the equation of motion. The equation is further written in a matrix form" + ] + }, + { + "image_filename": "designv11_92_0002639_chicc.2018.8483719-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002639_chicc.2018.8483719-Figure1-1.png", + "caption": "Fig. 1: Body-fixed frame and earth-fixed frame for UAV", + "texts": [ + " Then, the system states x1 and x2 converge to zero, if the control law is designed as u = \u2212\u03b1sign(x2)\u2212 \u03b2sign(x1) (2) with the adaption strategy \u03c1\u0307 = \u2212\u03b3\u03c1+ c(|uav|+ \u03c3) \u03c4 u\u0307av = \u2212uav + u (3) where \u03c1 can be considered as an estimation of the upper bound of the disturbance, uav is the average control, \u03c4 > 0 is the average filter time constant, \u03c3 > 0 is a constant which guarantees a desired minimum control level for start-up, \u03b1 = \u03c1, \u03b2 = 2\u03c1, and c > \u03b3 > 0. The configuration of the quadrotor UAV is given in Fig.1. It is composed by a rigid cross frame and four rotors. The position(x,y and z direction) and attitude(roll,pitch and yaw) motion can be achieved through an appropriate combination of the rotors 1 to 4. The 3DOF rotational motions of a quadrotor UAV is described by \u0398\u0307 = W\u03a9 I\u03a9\u0307 = \u2212\u03a9\u00d7 I\u03a9+ \u03c4 +\u0394(t) (4) where \u0398 = [\u03c6, \u03b8, \u03c8]T is the Euler angle, \u03a9 = [\u03c9x, \u03c9y, \u03c9z] T is the attitude angular velocity, \u0394 represents the external disturbances, \u03c4 = [\u03c41, \u03c42, \u03c43] T \u2208 R3 is the control torque vector, and I = diag[Ix, Iy, Iz] is a symmetric positive definite constant matrix" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003222_ciced.2018.8592275-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003222_ciced.2018.8592275-Figure4-1.png", + "caption": "Fig. 4 The section of the earthing ring does not meet the requirements 3) In order to satisfy the puncture performance of insulated", + "texts": [ + " And the voltage level of the 10kV earthing ring can always meet the requirements of the voltage strength of the 1kV insulated wire. In fact; the mechanical puncture force should be taken into account as well as the insulation electrical strength when selecting the earthing ring. It is stipulated in the technical requirements of the earthing ring that after the insulated wire is punctured, its running tensile force shall not be less than 95% of the calculated tensile force. The section of the earthing ring does not meet the requirements are shown in Figure 4. wire with different voltage levels, contact blade's tooth length, radian, shape, hardness and shearing torque of tensional shearing bolt (or nut) in the design of earthing rings needs to be adjusted many times, it is not feasible until the optimal electrical and mechanical performance requirements are met. Therefore, the 10kV test earthing ring, which is used with 10kV insulated wire, in addition to When lightning over voltage is flash-over, the instantaneous lightning current is very large\uff0c but the time is very short, it will only form a breakdown hole in the overhead insulated wire and will not burn the wire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001252_jctn.2008.828-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001252_jctn.2008.828-Figure3-1.png", + "caption": "Fig. 3. The FEM of temperature field of multi-layers thin-wall formed by DMLS.", + "texts": [ + " Based on the conversation of energy theory, the matrix expression of instantaneous heat exchange is expressed as Eq. (2). C T T + K T T = Q T (2) Where [K T ] is conduction matrix, which includes conduction coefficient, convection coefficient, radiance coefficient; [C T ] is specific heat matrix; {T } is temperature vector of the node; {T } is the temperature differential coefficient to time; [Q T ] is heat flow ratio vector of the node, and it includes the created heat energy. The calculating model of FE. The FEM of temperature field of multi-layers thin-wall formed by DMLS is expressed in Figure 3. The bottom side of the model is substrate with 9\u00d7 1 8\u00d7 0 9 mm, the upside is thin-wall part with 7 5\u00d70 45\u00d71 2 mm. To express the distribution of temperature correctly, the meshing of sintering area is smaller, and the rest part is larger. In the calculation, Solid 70 element with six sides and eight nodes is adopted. The cell length of work area and vicinity is 0 15\u00d7 0 15\u00d7 0 15 mm, and the rest is 0 3\u00d7 0 15\u00d70 3 mm. The substrate is 45 steel, and the sintering material is N -based alloy. The sintering part is divided into eight layers, the thickness of every layer is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002221_978-3-319-99270-9_27-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002221_978-3-319-99270-9_27-Figure6-1.png", + "caption": "Fig. 6. The figure of the prototype for two-stage radial gas turbine under investigation.", + "texts": [ + " Furthermore, the center of mass of the rotor is close to the drive end or rotor (Bearing 2). Therefore, the force in the Bearing 2 is higher than the Bearing 1. After obtaining the contact force between the balls and inner race, the maximum Hertzian contact stress between balls and bearing ring is obtained (Eq. 10). The highest contact stress in the contact of the rotor and Bearing1 and Bearing 2 found to be 3451 and 3549 MPa, respectively. The prototype of two-stage radial gas turbine shown in Fig. 6 produces 400 kW electricity. The nominal operation speed of both rotors is 550 Hz. The dropdown test of the rotor was carried out at two different speeds: 60 and 100 Hz. Before the dropdown, the rotor was supported by active magnetic bearings. Then, the magnetic bearing source was switched off and the unit is shut down. The measurement setup is equipped with two non-contact displacement sensors at the location of each backup bearings and the relative displacement of the rotor is recorded. These sensors are mounted at 45\u00b0 with respect to the vertical axis of the rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure6-1.png", + "caption": "Fig. 6. Second metamorphic limb with phase e2", + "texts": [ + " The metamorphic limb is synthesized with the phase e1 in Fig. 5. Link a is fixed to base, and rotational joint of axis 1 is locked. Axis 2 rotate along with axis v2 direction. Axis 3 rotate along with axis w2 direction. Axis 4 and axis 5 are parallel to axis 3 with w2 direction, end output link translate along with axis 6. Then metamorphic limb takes on mobility five which can realize motion type of submanifold T (3) \u00b7U(p2,w2,v2). Keeping rotational joint of axis 1locked, rotate link c to axis 3 parallel with axis u2 direction in Fig. 6. Axis 2 rotate along with axis v2 direction, axis 4 and axis 5 are parallel to axis 3 with u2 direction. Then metamorphic limb satisfy form of R(q21,v2) and R(q22,u2) \u00b7 R(q23,u2) \u00b7 R(q24,u2) \u00b7 T (v2) which generated by T (3) \u00b7U(q2,u2,v2) that have mobility five and can realize motion type of such submanifold. 1) Type synthesis of metamorphic joint of third limb: T2(w3) \u00b7 S(p3) and T2(v3) \u00b7 S(q3) are submanifolds of third limb of motion type of 1R2T and 2R1T, respectively. We can find that motion joint of third limb should satisfy one condition that transform between T2(w3) \u00b7S(p3) and T2(v3) \u00b7 S(q3)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.11-1.png", + "caption": "FIGURE 16.11 (a) Butt joint and (b) lap joint specifications in laser welding.", + "texts": [ + " Fluctuation in the energy absorbed may also be minimized by superimposing a pulsed beam over the CW welding beam. This is referred to as the rippled mode laser output or hyperpulse (see Section 14.1.4). The pulse rapidly melts the material surface, thereby increasing its absorption coefficient to a high enough value that the CW beam is efficiently and uniformly absorbed. The joint fitting requirements for lap joints are not as critical as for butt joints since in the lap joint, the laser beam is made to penetrate through one piece into the underlying component (Fig. 16.11). Lap joints are more appropriately used when the upper component is relatively thin (typically less than 1 mm in thickness). Thus far, a significant number of laser welds that are made are autogenous, involving no filler metal. However, filler metal may be used where necessary, especially for nonsquare butt joints or when the butt gap is relatively large. Processing efficiency is discussed in general terms in Section 14.2. For laser welding, the overall efficiency, \u03b7, may be expressed as \u03b7 = power used in melting solid power input to the laser generator (16" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002874_icstcc.2018.8540682-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002874_icstcc.2018.8540682-Figure1-1.png", + "caption": "Fig. 1 Fixed-wing aircraft and Translational and Rotational Forces and Speeds.", + "texts": [ + " The fixed-wing UAV is a significantly complicated system to be modelled considering all of the external forces and torques that also depend on the shape of the vehicle [8]. However, over the years under some assumptions which is discussed in this section that have proved to be valid have led to the model that is used in this paper [9]. An aircraft has conventionally 6 degrees of freedom dynamics, 3 torques and 3 forces govern the motion of an airplane. Due to aerodynamic and gravitational forces and torques, there are 3 translational and 3 angular velocities. This is illustrated in the Fig. 1[10]. 978-1-5386-4444-7/18/$31.00 \u00a92018 IEEE 740 \ud835\udc62, \ud835\udc63, \ud835\udc64 in Fig.1 are the translational velocities in the body reference frame. Additionally, if the effect of the wind is omitted, these speeds are equal to their earth reference frame counterparts. The roll, pitch and yaw angular speeds are denoted by \ud835\udc5d, \ud835\udc5e, \ud835\udc5f respectively. \ud835\udc4b, \ud835\udc4c, \ud835\udc4d denote translational forces and \ud835\udc3e, \ud835\udc3f, \ud835\udc40 represents angular forces that act on the aircraft in the axes as illustrated in the Fig. 1. Under the flat non-rotating earth assumptions, for a single rigid body that is symmetric about \ud835\udc67\ud835\udc65 plane the equations that describe the motion are given as, \ud835\udc4b \u2212 \ud835\udc5a\ud835\udc54sin\ud835\udf03 = \ud835\udc5a(?\u0307? + \ud835\udc5e\ud835\udc64 \u2212 \ud835\udc5f\ud835\udc63) (1) \ud835\udc4c + \ud835\udc5a\ud835\udc54cos\ud835\udf03sin\ud835\udf19 = \ud835\udc5a(?\u0307? + \ud835\udc5f\ud835\udc62 \u2212 \ud835\udc5d\ud835\udc64) (2) \ud835\udc4d + \ud835\udc5a\ud835\udc54cos\ud835\udf03cos\ud835\udf19 = \ud835\udc5a(?\u0307? + \ud835\udc5d\ud835\udc63 \u2212 \ud835\udc5e\ud835\udc62) (3) \ud835\udc3f = \ud835\udc3c\ud835\udc65?\u0307? \u2212 \ud835\udc3c\ud835\udc67\ud835\udc65?\u0307? + \ud835\udc5e\ud835\udc5f(\ud835\udc3c\ud835\udc67 \u2212 \ud835\udc3c\ud835\udc66) \u2212 \ud835\udc3c\ud835\udc67\ud835\udc65\ud835\udc5d\ud835\udc5e (4) \ud835\udc40 = \ud835\udc3c\ud835\udc66?\u0307? + \ud835\udc5f\ud835\udc5d(\ud835\udc3c\ud835\udc65 \u2212 \ud835\udc3c\ud835\udc67) + \ud835\udc3c\ud835\udc67\ud835\udc65(\ud835\udc5d2 \u2212 \ud835\udc5f2) (5) \ud835\udc41 = \ud835\udc3c\ud835\udc67?\u0307? \u2212 \ud835\udc3c\ud835\udc67\ud835\udc65?\u0307? + \ud835\udc5d\ud835\udc5e(\ud835\udc3c\ud835\udc66 \u2212 \ud835\udc3c\ud835\udc65) + \ud835\udc3c\ud835\udc67\ud835\udc65\ud835\udc5e\ud835\udc5f (6) \ud835\udc5d = ?\u0307? \u2212 ?\u0307?sin\ud835\udf03 (7) \ud835\udc5e = ?\u0307?cos\ud835\udf19 + ?\u0307?cos\ud835\udf03sin\ud835\udf19 (8) \ud835\udc5f = ?\u0307?cos\ud835\udf03cos\ud835\udf19 \u2212 ?\u0307?sin\ud835\udf19 (9) ?\u0307? = \ud835\udc5d + (\ud835\udc5esin\ud835\udf19 + \ud835\udc5fcos\ud835\udf19)tan\ud835\udf03 (10) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001101_iros.2008.4651186-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001101_iros.2008.4651186-Figure3-1.png", + "caption": "Fig. 3. Force balance on slope traversing case", + "texts": [ + " From the figure, it is easily found that the lateral velocity of vehicle must be greater than or equal to zero for successful slope traversal. An analysis of a slope climbing of a rover is simply determined by a uniaxial force balance between drawbar pull (longitudinal tractive force) of wheel and a vehicle\u2019s traction load due to the gravity. However, when the rover traverses a sandy slope, the wheel generates not only longitudinal force but also lateral force since the rover makes steering maneuvers to traverse a slope. Fig. 3 describes a schematic figure of force balance on slope traversing case. Focusing on the force balance between the tractive force generated by wheels and vehicle traction load, the tractive force must be larger than the traction load in order to achieve its slope traversing. Further, a sideslip of wheel can be observed in slope traversing. This slippage into lateral direction of wheel is measured by slip angle \u03b2. The detail of the slip angle is explained in Section III, later. Because of this sideslip, the wheel generates lateral force named Side force Fy ", + " It can be clearly seen that tractive force of vehicle is composed of a summation from Fc1 to Fc4, which are cornering forces of each wheel. Then, the cornering force Fc is given as a function of drawbar pull Fx and side force Fy : Fc = Fx sin \u03b2 + Fy cos \u03b2 (1) From the above, the rover can traverse when a summation of cornering forces overcomes the traction load W sin \u03b8s. Therefore, one of criteria for the slope traversability is determined as follows: (Fc1 + Fc2 + Fc3 + Fc4) \u2265 W sin \u03b8s (2) On the other hand, another wheel force, FT , can be composed as seen in Fig. 3. In this research, FT is named as Thrust force consisting of drawbar pull and side force: FT = Fx cos \u03b2 \u2212 Fy sin\u03b2 (3) The thrust force needs to be greater than zero so that the rover can travel the slope ahead. Then, another criteria for the traversability is simply defined as: FT1 = FT2 = FT3 = FT4 \u2265 0 (4) From the above note, there is an equilibrium point where the traction load is equal to a summation of cornering forces. Then, it is considered that the rover can traverse a slope along with straight line at this force equilibrium point", + " The trafficability analysis using this diagram is as follows: First, a cornering force of each wheel which balances with a traction load on an arbitrary slope angle can be determined using (7). For instance, a traction load is assumed to be 3.0 [N]. Then, this turns out that the required cornering force to traverse the slope is given as Fc \u2265 3.0. Also, referring to another criteria for slope traversing capability determined in (4), the required thrust force is given as FT \u2265 0. Subsequently, drawing these two boundary lines (Fc \u2265 3.0 and FT \u2265 0) on the diagram, it can be seen that the diagram is divided into two regions, namely traversable/untraversable regions. Here, as shown in Fig. 3, the steering angle is assumed to be equivalent to the slip angle if the rover successfully traverses the slope in straight ahead. Therefore, the thrustcornering characteristic diagram determines the trafficability of the rover based on given steering angles: as shown in Fig. 8, it is expected that the rover having more than 12 [deg] of steering angles can traverse and climb the corresponding slope since the characteristics curve of \u03b2 = 12 [deg] is larger than the required cornering force and is included in the traversable region" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.28-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.28-1.png", + "caption": "Fig. 11.28 Relation between the fore\u2013aft force and side force and the relation between the fore\u2013aft force and self-aligning torque. Reproduced fromRef. [2] with the permission ofGuranpuri-Shuppan", + "texts": [ + "104) is given by the condition tan2 a\u00fe s2 3lsFz=CFa\u00f0 \u00de2: \u00f011:106\u00de 13Problem 11.3. 14See Footnote 13, Note 11.11. Figure 11.26 shows the calculation of the fore\u2013aft force for various slip angles and slip ratios. Figure 11.27 shows the calculation of the side force and self-aligning torque for various slip angles and slip ratios. Combining the results of Figs. 11.26 and 11.27, the relation between the fore\u2013aft force and side force and the relation between the fore\u2013aft force and self-aligning torque are obtained as shown in Fig. 11.28. The calculations use parameters of the load (Fz = 4 kN), contact length (l = 243 mm), lateral spring rate of the tire (Ky = 2.5 kN/cm), braking and driving stiffness (CFs = CFa = 57.2 kN) and friction coefficient (ls = 1.0, ld = 0.7). According to Sakai [2], the results in Figs. 11.26, 11.27 and 11.28 agree well with the measurements of bias tires but not those of radial tires. A comparison of Eq. (11.103) with Eq. (11.104) reveals that the side force under the braking condition is stronger than that under the driving condition if the slip ratio and slip angle are small" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002913_1.5080071-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002913_1.5080071-Figure1-1.png", + "caption": "FIGURE 1. Components of wind turbine drive train", + "texts": [ + " The lumped mass dynamic model is developed to get the vibration response of wind turbine drive train. The dynamic model accounts for time varying gear mesh stiffness, dynamic transmission error and bearing elasticity. The natural frequencies and corresponding mode shapes are obtained using state space approach method. The 22 equations of motion obtained are solved for system responses by Newmark time-step integration algorithm. The vibration displacement and acceleration are obtained in time and frequency domain. MODELLING OF DYNAMICS OF DRIVE TRAIN Figure 1 shows major components of a wind turbine drive train. A two stage planetary with each stage consisting of three planets, sun and ring fixed and a single stage parallel gear box is typically considered here. All the gears assumed are to be spur. Rotor is the input, where the driving torque is applied and then subsequently transferred to the planet carrier of first planetary stage which rotates the planets that are symmetrically positioned at an angle. Planets then rotate the sun gears which are in mesh with sun" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure16-1.png", + "caption": "FIG. 16: UNROLLED GEOMETRY FOR A CIRCLE (THIN GREEN LINE) AND THE SQUARE FLANGE (THICK BLUE LINE).", + "texts": [ + " In concept, the rotary positioning is aligned with planar deposition solutions, and the rotary related cases are distinct, such as the decisions with respect to the mandrel / build cylinder, the creation of geometry to support roughing and finishing operations, and build strategies to address nonsymmetric rotary components. AM rotary tool path support provides designers a process planning strategy that could avoid the introduction of support material as \u2018out of plane\u2019 features can be built up incrementally. Ideally, the AM rotary tool paths could support nonrotational features. One solution for set for the examples presented here can be achieved through unroll-roll functions for the feature edge geometry. Consider the flange in Fig. 16. When unrolling the bounding swept geometry (a circle), a straight line results; whereas, the unrolled edge geometry for the square flange appears as a set of semi-circles, where the farthest protrusion points for the flange touch the straight line. The unroll operation is performed for the cam example (Fig. 17). Offsetting the upper bound by the desired slicing layer height, and trimming these offset curves to the unrolled geometry edges will generate the stop-start points for the deposition segments for each deposition layer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003819_b978-0-12-812939-5.00002-1-Figure2.5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003819_b978-0-12-812939-5.00002-1-Figure2.5-1.png", + "caption": "FIG. 2.5 Template with added foot segment.", + "texts": [ + "168) FyEE\u2217 + X2 j\u00bc1 \u03bbj t\u00f0 \u00de \u2202gj \u2202y1 \u00bcFyEE\u2217 + \u03bb2 t\u00f0 \u00de (2.169) m1 +m2\u00f0 \u00de\u20acyCM + m1 +m2\u00f0 \u00deg \u00bcFyEE\u2217 + \u03bb2 t\u00f0 \u00de (2.170) provide nice interpretation for Lagrange multipliers. Namely, the forces of constraints are exactly ground reaction forces we obtained earlier \u03bb1 t\u00f0 \u00de\u00bc FgrX (2.171) \u03bb2 t\u00f0 \u00de\u00bcFgrY (2.172) 2.3.9 Locomotor Dynamics, Ground Reaction Forces, and ZMP Imagine adding additional segment that we could interpret as \u201cfoot\u201d attached to O1 joint of our template system. The foot is assumed flat on the ground and subject to ground reaction forces (Fig. 2.5). Eq. (2.76) for joint 1, formulated within the Newton-Euler method, is now modified to \u03c4 ! 1 + r ! CM w=foot r ! O1 m mfoot g ! + r ! EE r ! O1 F ! EE\u2217 \u00bc d L ! 1 O1\u00f0 \u00de + L ! 2 O1\u00f0 \u00de dt (2.173) with r ! CM w=foot being CM of links 1 and 2 and mfoot being the foot mass. Note that there is no moment due to ground reaction forces in Eq. (2.173). If we added the term that would be an erroneous example of \u201cdouble counting,\u201d as torques applied to adjacent links are equal in magnitude and opposite in direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003415_phm-chongqing.2018.00022-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003415_phm-chongqing.2018.00022-Figure2-1.png", + "caption": "Fig. 2 Gear tooth spalling with spherical shape", + "texts": [ + " The transition curve can be expressed as a function of an angle parameter \u03b3 as [14] ( ) ( ) ( ) ( ) ( ) ( ) cos sin sin sin cos sin t p tr xt xt p tr A x R R A y h R R \u03c6 \u03b3 \u03c6 \u03b3 \u03c6 \u03b3 \u03c6 \u03b3 = \u2212 + \u2212 = = \u2212 + \u2212 (2) Where 0 / 2\u03b1 \u03b3 \u03c0\u2264 \u2264 , pR is the radius of pitch circle, ( )0/ 1 sin( )trR c m \u03b1\u2217= \u2212 , ( ) trA h c m R\u2217 \u2217= + \u00d7 \u2212 , ( )/ tan( ) / pA B R\u03c6 \u03b3= + , in which * 0 0/ 4 tan( ) cos( )a trB m h m R\u03c0 \u03b1 \u03b1= + + . For standard spur gears 1h\u2217 = , 0.25c\u2217 = , and 0 20\u03b1 = . III. THE PROPOSED SPHERICAL SHAPE-BASED METHOD ON Fig. 2 and Fig. 3 show the proposed spherical shape-based method for modelling the curve-bottom feature of a gear tooth spall as observed in practice. The tooth spall is constructed by removing the intersection between the spherical surface and gear tooth. The crosssection view cuts through the center of the sphere, see Fig. 2. _sk stx is the start position of the curved spall, _ maxskh is the maximum depth, sk\u03b8 is the incline angle with respect to the x axis. skO is the geometric center of the spherical, skr is the corresponding radius. The subscript k denotes the kth tooth spall. The cross-section model of an arbitrary position of the tooth spall is shown in Fig. 3. For a spherical tooth spall in practice, _sk stx , _ maxskw , sk\u03b8 and _ maxskh are determined by the severity of the tooth spall. With knowledge of these parameters, the radius of the spherical can be can be defined as: ( )2 2 _ max _ max _ max _ max 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003622_icsee.2018.8645997-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003622_icsee.2018.8645997-Figure2-1.png", + "caption": "Fig. 2: The three basic settings of additional sensors to the micro drone: in all models an optical-flow sensor was installed at the bottom of the drone and range sensors were placed at the front / sides / back of the drone. Left: the drone with a single front Time of Flight ranging sensor (ToF ) (see Figure 5 - up). Middle: the drone with five ToF sensors facing front, left, back, right and up (see Figure 5 - down). Right: three ToF sensors facing front and +-45 degrees (front left and front right).", + "texts": [ + " Larger drones of 60 grams and above can have a real companion computer - which is usually based in embedded linux (see the Tello drone by Ryze Tech) yet, at this weight the drone size is about 15-20 centimeters in diameter which makes it considerably less suitable for indoor autonomous flying (i.e., big birds can not fly in narrow caves - while bets and insects do - easily). In this work we have mainly focused on the Crazyfile 2.0 drone which is a 27 gram open source drone which is commonly used by researchers for AI and BioInspired applications. Several sensor settings were examined, see Figure 2 for the main three sensor settings. High-end commercial drones are often equipped obstacle avoiding sensors (see for example DJI\u2019s Phantom4 pro drone, or Parrot\u2019s SLAM dunk), such sensors uses high computing power, yet the are only able to detect relatively solid obstacles and require proper light conditions. In fact, such drones do not have generic capability for performing indoor autonomous mapping. Moreover, in natural scenarios which may include flying in a forest or at night (poor light conditions) commercial drones are still unable to perform an autonomous flights" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001038_iemdc.2009.5075448-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001038_iemdc.2009.5075448-Figure3-1.png", + "caption": "Figure 3. The stationary reference frame and the stator flux oriented reference frames.", + "texts": [ + " The radius of propeller for 2 [kW] wound rotor induction generator can be obtained from calculation of torque coefficient of propeller typed windmill in reference [6]. For example, when the wind speed is set to be 12 [m/s] and the output power is set to be 2 [kW], the rotational speed is N = 820 [rpm] from Figure 2. In this case, the tip speed ratio \u03bb and torque coefficient Ct is approximately 0.05 from Figure 1. 1808978-1-4244-4252-2/09/$25.00 \u00a92009 IEEE The vector control is applied to control the active and reactive power of a wound rotor induction generator. Figure 3 shows a reference frame used in this study [7]. The stator coordinate (stationary coordinate) is the \u03b1-\u03b2 coordinate, and the stator flux oriented coordinate is the d-q coordinate. Figure 4 shows the overall system configuration [5]. A threephase stator voltage and stator current are converted to the stator flux oriented coordinate in order to calculate active and reactive power. The real and reference values for these are compared and made to pass the PI control parts to obtain the rotor current reference value in the d-q coordinate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001065_icmtma.2009.276-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001065_icmtma.2009.276-Figure3-1.png", + "caption": "Figure 3. Model of substructure", + "texts": [ + " DETAILS OF THE COMPUTATIONAL MODEL The compressor impeller is made up of identical segments in the circumferential direction and hence constitutes a rotationally periodic structure. The impeller with a shaft hole 34mm in diameter is mounted onto the sleeve and the compressor shaft via interference fit. The material of the sleeve and shaft is steel and the material of the impeller is aluminum. The computation needs to be carried out by using 3D model, instead of by axial symmetric one. The computational model of substructure and the impeller is shown in Figs. 2. The model of substructure is shown in Fig 3. The main loads applied to the compressor are interference-fit load and centrifugal forces caused by highspeed rotation when operating. 1\u03b4 is defined as the amount of interference between impeller and the shaft sleeve and 2\u03b4 as the amount of interference between the shaft sleeve and shaft. For different values of 1\u03b4 =0.04 mm and 2\u03b4 =0.05mm, the load cases are computed respectively for various rotational speed n (0, 24000, 25000, 27000, 29000rpm), various coefficients of friction \u03bc (0.1, 0.15, 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003692_s1068799818040104-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003692_s1068799818040104-Figure4-1.png", + "caption": "Fig. 4. The ring subject to a radial force and distributed tangent tractions.", + "texts": [ + " A given deformed shape of the ring emerged from a pointwise radial force jQ can be decomposed into the Fourier series as =0 ( ) = cos( ),j j k k w Q k \u221e \u03d5 \u03bb \u03d5\u2211 (10) where multipliers 0 1 = ( ) if = 0 2 w d k \u03c0 \u2212\u03c0 \u03bb \u03d5 \u03d5 \u03c0 \u222b ; (11) 1 = ( )cos( ) if > 0k w k d k \u03c0 \u2212\u03c0 \u03bb \u03d5 \u03d5 \u03d5 \u03c0 \u222b (12) are physically nothing but the compliance coefficients. According to the superposition principle valid for small deformations, deflection of the ring mid-surface resultant from N radial forces jQ acting at an angle j\u03c8 can be expressed as ( ) =1 =1 =0 ( ) = ( ) = cos . N N j j k j j j k w w Q k \u221e \u03d5 \u03d5 \u03bb \u03c8 \u2212 \u03d5\u2211 \u2211 \u2211 (13) In order to find factors k\u03bb , let us recall the analytical solution of [14] for the ring subject to a pointwise radial force, applied at an arbitrary point and equilibrated by distributed tractions as shown in Fig. 4: 3 ( ) = ( ), 4 QR w EI \u03d5 \u03c7 \u03d5 \u03c0 (14) where 2 23 ( ) = ( )sin cos 2 . 2 4 3 \u239b \u239e\u239b \u239e\u03d5 \u03c0 \u03c7 \u03d5 \u2212 \u03d5 \u2212 \u03c0 \u03d5 + \u03c0\u03d5 \u2212 + \u2212 \u03d5 +\u239c \u239f\u239c \u239f\u239c \u239f\u239d \u23a0\u239d \u23a0 (15) Introducing 3 = 4k k R EI \u03bb \u03b3 \u03c0 , one may write the first five coefficients of the Fourier series expansion for the function ( )\u03c7 \u03d5 : IVANNIKOV et al. RUSSIAN AERONAUTICS Vol. 61 No. 4 2018 572 0 1 2 3 4 4 1 4 = 0; = 0; = ; = ; = . 9 16 225 \u03b3 \u03b3 \u03b3 \u03b3 \u03b3 (16) Solution (14) does not take into account the kinematic boundary conditions imposed to the ring by the support, which in real situations may supplement additional, sometimes quite substantial, stiffness. Due to this fact, it is recommended to use a 3D model of the support\u2013bearing assembly (of course, in the case of its availability) to precise the k\u03bb values. To this end the corresponding three-dimensional spatial model is loaded by a unitary radial force at an arbitrary radial point as shown in Fig. 4. The radial deflections obtained are then used for numerical integration of expressions (11). The constructed approximation has some disadvantages. First, the tensile ring deformations are omitted. Second, if the profile of the acting radial forces tends to sin or cosin, i.e. ( ) = cos( ),o aQ Q Q\u03c8 + \u03c8 (17) the resultant radial displacements ( ) 0w \u03d5 \u2192 . It is easy to establish this fact. Due to orthogonality of the individual members of expansion (10), the following terms, entering (12), vanish: ( )( ) ( ) =1 cos cos = 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000118_20080706-5-kr-1001.00015-FigureA.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000118_20080706-5-kr-1001.00015-FigureA.1-1.png", + "caption": "Fig. A.1. slip angles of front and rear tires", + "texts": [ + "5) where m is the mass of the automobile, I is the moment of inertia, M is the total torque of the center of gravity, \u03c8\u0307 is the yaw rate and \u03b2 is the angle between the orientation of the automobile and the velocity vector (Fig. 4). Since ve\u03042 = 0 holds, ve\u03042 and v\u0307e\u03042 of (A.3) and (A.4) are zero, respectively. For simplicity, we describe ve\u03041 as v. In this paper, fe\u03041 is assumed to be given by damping force \u00b5(v \u2212 v0) and driving force w. Then (A.3) is written by the following equation: mv\u0307(t) = \u00b5(v(t) \u2212 v0) + w(t) (A.6) wherev0 is the operating point of the velocity. Moreover, we assume the side forces are generated due to the slip angles of each tire(See Fig. A.1, such assumption is also used in Abe [1979]). The each forces are given by Kf\u03b2f (t)e\u03042(t) and Kr\u03b2r(t)e\u03042(t) where Kf and Kr are the cornering powers of the front and the rear tires respectively, and \u03b2f (t) and \u03b2r(t) are the slip angles (angle from orientation of each tire to velocity vector). Furthermore, \u03b2f (t) and \u03b2r(t) can be approximated by \u03b2f (t) = \u03b2(t) + lf \u03c8\u0307(t) v(t) \u2212 \u03b4(t), (A.7) \u03b2r(t) = \u03b2(t) \u2212 lr\u03c8\u0307(t) v(t) , (A.8) where lf and lr are the distances from the center of gravity to the positions of the front and the rear tires" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001996_icasi.2018.8394258-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001996_icasi.2018.8394258-Figure2-1.png", + "caption": "Fig. 2 Reference frames of the underwater vehicle.", + "texts": [ + " In order to avoid the singularity problem of Euler angles and better describe large angle attitude maneuver of the CMG-actuated underwater vehicle, the unit quaternion [11] is given as follows 1 2 3[ , , , ] [ , ] 2 2 T Ts c q (1) where 1T q q , c=:cos , =:sins . and are the rotation axis vector and rotation angle, respectively. Thus the transformation relationship related to the linear velocity between the earth-fixed frame and body-fixed frame can be expressed in the form 1 1( )E q (2) where 1 [ , , ]Tx y z and 1 [ , , ]Tu v w denote the position vector and linear velocity vector of the underwater vehicle respectively. Additionally, Denote the Euler angles Pitch , Roll- and Yaw- (see Fig. 2), rotation matrix ( )E q and the relationship between Euler angles are given by 2 2 2 3 1 2 3 1 3 2 2 2 1 2 3 1 3 2 2 2 3 3 2 11 1 2 1 3 2 1 2( ) 2( ) 2( ) ( ) 2( ) 1 2( ) 2( ) 2 2( ) 1 2( ) E \uff08 - \uff09 q (3) 1 2 3 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 c s c s c s c c s s s c c c s s s c c c c s s s q (4) Consider that motions of the vehicle are always subject to the effects of hydrodynamic forces and moments, and thus the center of gravity and buoyancy are assumed to be the same level for simplicity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001132_2009-01-2105-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001132_2009-01-2105-Figure9-1.png", + "caption": "Figure 9 shows the setup for the experimental modal analysis on the loaded tyre. The radial, tangential and axial response is measured in each point of the grid with a triaxial accelerometer (PCB 356A15). In order to avoid mass loading of the structure, only 6 accelerometers are used simultaneously and they are distributed evenly over the tyre circumference. Comparison between the measured and predicted natural frequencies (figure 10) shows that the predicted natural frequencies are within 5% of the measured natural frequencies.", + "texts": [ + " EIGENVALUE ANALYSIS OF LOADED TYRE \u2013 After the static tyre loading, an eigenvalue analysis is performed on the inflated, deformed tyre which is clamped at the wheel centre. The nodes which are in contact with the road surface remain fixed during the eigenvalue calculation. Figure 8 shows the split of a double pole of the unloaded tyre into two single poles as a result of the tyre loading. Appendix A explains the naming of the loaded tyre modes. An experimental modal analysis is performed on the loaded tyre. The wheel is clamped at a rigid spindle and loaded such that a deformation of 16 mm is obtained. Fig.9: Test setup for the experimental modal analysis on the loaded tyre. electrodynamic shaker fixed clamp LOADED TYRE FREQUENCY RESPONSE FUNCTIONS \u2013 Similar to the unloaded tyre, the steadystate linearized response of the loaded tyre-wheel model to a harmonic excitation at the treadband is calculated up to 300 Hz. Figure 11 shows the comparison between the calculated and measured response in radial, tangential and axial direction due to a harmonic force. The excitation and response position is the same as in the test on the unloaded tyre and is shown above the graphs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001617_s40799-018-0232-7-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001617_s40799-018-0232-7-Figure1-1.png", + "caption": "Fig. 1 Schematic of the testing bench loaded at the end and with a strain gage at a known distance (d = 50 mm)", + "texts": [ + " During the experimental stage, the use of a type T strain gage connected to a data acquisition system is proposed, where the attachment, signal preparation and setup of the interface are essential parts of the lecture. Finally, the results obtained by the above three methods are compared in order to discuss the differences observed and possible sources thereof. Once the activities are fulfilled, students highlights their results and conclusions in written reports, some of them are listed in the results section. The following section describes in detail the issue addressed in this work, as well as the materials and methods used. This work proposes an experimental testbed whose schematic is shown in Fig. 1. This testing bench consists of a metallic frame which supports a cantilever circular hollow bar. Additionally, a solid cylinder, where the load is applied, is held at 90 degrees with respect to the hollow bar. Both components are made of AISI 1018 steel, the mechanical properties of which are shown in Table 1. The measurement system is set up by using various strain gages with different connection diagrams, which have been used to analyse other cases reported in the literature [25, 26]. The configuration proposed and analysed in this work uses a type T strain gage for measuring normal strains in two orthogonal directions on the solid cylinder", + " The initial results are the longitudinal and tangential strains, at the defined area which represents the attached strain gage. For this purpose, a local cylindrical coordinate system should be used on the solid cylinder. It is suggested that the students iterate different mesh sizes, as well as assessing other results of the structure such as the normal stress throughout the length of the cylinder. Once the numerical model has been defined, the analytical stage is carried out to determine strains in the system as is detailed in the next section. The effect of a given proposed load on the solid cylinder, as shown in Fig. 1, can be defined by considering the cylinder as a cantilever beam, as seen in Fig. 4. Once the dimensions of the beam and the system are defined, the maximum stress at a point of interest can be determined through equation (1), which involves stress and bending moment for small displacements [27, 28]. \u03c3l \u00bc Mc I \u00f01\u00de According to equation (1), for a given component with circular section, c and inertial moment I depend on the cross sectional radius r. The bending moment (M) is caused by the applied load and the distance to a given point of interest at a distance (d)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001169_dscc2009-2731-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001169_dscc2009-2731-Figure2-1.png", + "caption": "FIGURE 2. A TORSIONAL SYSTEM WITH LUMPED INERTIAS AND TORSIONAL ELASTIC SPRINGS.", + "texts": [ + " The values of the loss factor associated with the coupling and the engine are supplied by the manufacturers as 0.124 and 0.08, respectively. A relatively small value of 0.0015 is assumed for rest of the driveline components. The equations of motion (EOMs) of the test cell system are derived by representing it as multiple lumped inertias, 1 2 , , , ,\u22ef n J J J connected with torsional elastic springs with stiffnesses, 1 2 1 , , , . \u2212 \u22ef n K K K The schematics of such a pure torsional system with n degree-of-freedom (DOFs) or inertias and n-1 torsional springs are shown in Fig. 2. Equating the equal and opposite torques at each mass, n distinct EOMs are formed as: ( ) ( ) ( ) ( ) 1 1 1 1 2 2 2 1 1 2 2 2 3 1 1 ( ) . n n n n n J K t J K K J K \u03b8 \u03b8 \u03b8 \u03c4 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u2212 \u2212 = \u2212 \u2212 + = \u2212 \u2212 \u2212 = \u2212 \u027a\u027a \u027a\u027a \u22ee \u027a\u027a (3) These EOMs can be expressed in a matrix form by using the rotational accelerations 1 2 , T n\u03b8 \u03b8 \u03b8 \u027a\u027a \u027a\u027a \u027a\u027a\u22ef the rotational displacements, [ ]1 2 , T n \u03b8 \u03b8 \u03b8\u22ef and the rotational inertia and stiffness matrices as: [ ]{ } [ ]{ } ( )J K T t\u03b8 \u03b8+ = \u027a\u027a (4) where, 2 Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000069_tee.20444-Figure17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000069_tee.20444-Figure17-1.png", + "caption": "Fig. 17 Trajectory with dynamic obstacle", + "texts": [ + "874) in turn by online learning to avoid the obstacle until it achieves the task finally. The run time until reaching the final goal is 65 s. The running trajectory of vehicle car is shown in Fig. 16. 4.3. With moving obstacle The initial position of moving obstacle car is set as (\u22128 m, 6 m, 0). By predicting and evaluating the future state of obstacle car and itself, vehicle selects (10 m, 6 m, 0.5\u03c0) (\u00b5 = 0.556) as sub-target firstly, then selects (2 m, 8 m, \u22120.75\u03c0) (\u00b5 = 0.849) to obtain the collision-free and low-cost path. The running trajectory of it is shown in Fig. 17. The time of reaching the final target is 62 s. 4.4. With static and moving obstacles The initial position of moving obstacle car is set as (\u22128 m, 12 m, 0). First, vehicle selects (10 m, 6 m, 0.5\u03c0) (\u00b5 = 0.556) and (8 m, 10 m, 0.75\u03c0) (\u00b5 = 0.644) as sub-target to evade the static obstacle and approach to the final target. But before it reaches (8 m, 10 m, 0.75\u03c0), it detects the moving obstacle car and had to reverse to guarantee the safety by selecting (10 m, 2 m, 0.5\u03c0 ) (\u00b5 = 0.668) and (8 m, 0 m, 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003519_rcar.2018.8621793-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003519_rcar.2018.8621793-Figure1-1.png", + "caption": "Figure 1. Coordinate frame of CSM", + "texts": [ + " The rest of this section will give a more detailed presentation and more in-depth analysis. The method of Camera Space Manipulation was proposed by Skaar in [18]. The method achieves robust and accurate visual guidance without the need to pre-calibrate cameras and robots, in addition, CSM does not require fast image processing required as visual serving methods. CSM establishes the mathematical relationship between robot internal joint angles and 2-D image plane coordinates in camera space which is called Camera Space Kinematics. As show in Fig. 1. The coordinate frame O-XYZ represents the robot's local coordinate frame, o-xyz is the camera coordinate frame, the origin of which is at the camera's focus, ,c cx y represents the image plane coordinate frame, and the origin is located in the center of the image, and the two axes parallel to the camera coordinate system x, y axis respectively. The three-dimensional position in the robot's local coordinate system is ( , , )x y z , and its corresponding coordinates in the image plane is ( , )c cx y , then the camera space kinematics can be expressed as: 2 2 2 2 1 2 3 4 2 3 1 4 2 4 1 3 5 2 2 2 2 2 3 1 4 1 2 3 4 1 2 3 4 6 ( ) 2( ) 2( ) 2( ) ( ) 2( C ) c c x C C C C x C C C C y C C C C z C y C C C C x C C C C y C C C z C where 1 2 3 4 5 6, , , , ,C C C C C C is the visual parameters of CSM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000443_detc2009-86926-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000443_detc2009-86926-Figure2-1.png", + "caption": "Figure 2. Hyper plane classifying two classes small margin", + "texts": [ + "com Proceedings of the ASME 2009 Inter ational Design Engineering Technical Conferences & Computers and Information in Engine ring Conference IDETC/CIE 20 9 Aug st 30 - September 2, 20 9, San Diego, California, USA DETC2009- Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 2 Copyright \u00a9 2009 by ASME Pattern recognition and classification using SVM is described here in brief, a more detailed description can be find in Ref. [7-8]. A simple case of two classes is considered, which can be separated by a linear classifier. Figure 2 shows triangles and squares stand for these two classes of sample points, respectively. Hyper plane H is one of the separation planes that separate two classes. H1 and H2 (shown by dashed lines) are the planes those are parallel to H and pass through the sample points closest to H in these two classes. Margin is the distance between H1 and H2. The SVM tries to place a linear boundary between the two different classes H1 and H2, and orientate it in such way that the margin is maximized, which results in least generalization error" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003866_ijabe.v12i2.3227-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003866_ijabe.v12i2.3227-Figure1-1.png", + "caption": "Figure 1 Dynamic tire model based on HPSO-SVM", + "texts": [ + " A new idea for predicting tire force of agricultural vehicle was proposed. Considering that the working ground of agricultural vehicles is mostly soft road[19,20], where tire stiffness is much greater than soil stiffness, tire deformation is relatively not obvious. Therefore, tire modeling under soft road conditions usually assumes that tire is a rigid wheel and only considers the deformation of soil. The dynamic tire model based on HPSO-SVM is composed of two parts: the ground tire model and the deformable terrain force model, as shown in Figure 1. This paper pays close attention to the ground tire model during the deformable terrain force model has already been covered thoroughly[4,5,7,10]. According to the force of the tire, the ground tire model calculates the motion state of the tire, and transmits it to the deformable terrain force model, and then the deformable terrain force model dynamically calculates the tire-soil contact force, and finally transmits the contact force to the ground tire model. The HPSO-SVM model is used to dynamically describe tread pattern, wheel spine, tire sidewall elasticity and inflation pressure between ground tire model and deformable terrain force model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003102_ecce.2018.8558121-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003102_ecce.2018.8558121-Figure9-1.png", + "caption": "Fig. 9. Prototype stator, rotor shaft and assembled machine with fan blade.", + "texts": [], + "surrounding_texts": [ + "Figure 11 shows measured shaft vibration without the fan blade. In the upper figure, the rotational speed is immediately accelerated up to the reference speed of 6120 r/min because of no load. The shaft vibration is high, and then, the rotor shaft is touch-down because the rotational speed is close to the tilting resonance frequency of the backward whirl as shown in Fig. 5. Although we tried to accelerate the motor more than 6120 r/min, the rotor shaft could not pass through the critical speed because the vibration is increased." + ] + }, + { + "image_filename": "designv11_92_0003394_6.2019-1747-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003394_6.2019-1747-Figure5-1.png", + "caption": "Figure 5: Components of the experiment testbed set up.", + "texts": [ + " The opposite sense simulation required an additional step to bring the hinge pass the initial snap through phase. This was done by first pressing the shell flat with a rigid pin, and then removing the pin and continuing to the symmetric fold configurations. These steps are excluded from the data. The full range of symmetric fold angle data is acquired despite the pin by stepping through the fold angle constraints in reverse, from fully folded to fully deployed. A mechanical testbed is designed to configure and control the asymmetric displacements, and a diagram of this design is presented in Figure 5. Two ATI six-axis force/torque transducers are used at the reference frames on the hinge to directly measure the full force/torque profile. The transducers are calibrated for torque measurements of 500 N-mm with 1/16th N-mm resolution and forces of 50 N in plane and 70 N out of plane with 1/80th N resolution. These sensors are aligned with the hinge such that the measurement frame of the sensor is coincident and orthogonally aligned to the hinge reference frames A0 and A1. The data from these hinges are then transformed into the frame alignments defined in Figure 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002227_s00706-018-2202-2-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002227_s00706-018-2202-2-Figure1-1.png", + "caption": "Fig. 1 Illustrations of a the injection cell in the setup at injection position and b exploded view of injection cell. The cell is installed in the bottom part (2) of the EC\u2013CE\u2013MS device next to buffer reservoirs for CE separation. The separation capillary (1) is installed in the top part. The injection cell consists of a bottom piece (4) with electrode slot and a cover piece (3) with electrical contacts. A silicone sealing ring prevents leakage of the sample", + "texts": [ + " The cell geometry was adapted to the existing EC\u2013CE\u2013MS setup [33] to allow for the usage of thin-film electrodes without changing the injection unit. Instead of a SPE, the injection cell with integrated thin-film electrode was installed in the injection unit. Due to electrode and cell dimensions, small sample volumes of only 10 mm3 or lower were sufficient for EC\u2013CE\u2013MS measurements, which is especially advantageous if only limited amount of sample is available. A schematic illustration of the final injection cell prototype is shown in Fig. 1. Polyether ether ketone (PEEK), a highly chemical resistant and mechanically stable material, was used for fabrication of the cell body. To prevent leakage, a sealing ring was integrated at the bottom of the open cell chamber. Thus, spreading of droplets could be avoided. As commercially available O-ring materials were attacked by organic solvents, a custom silicone sealing ring with appropriate dimensions (inner diameter 2 mm, outer diameter 4 mm, thickness 1 mm) was prepared. The electrical contact to the implemented thin-film electrode was achieved via spring contact probes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001765_978-3-319-91262-2_61-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001765_978-3-319-91262-2_61-Figure6-1.png", + "caption": "Fig. 6. Test 4", + "texts": [ + " Figure 4 shows that the algorithm didn\u2019t converge given the value of k and the few number of iterations. A third test was carried out using the same scenario that test 1, and test 2. But this time using a convergence constant of k = 0.4, and 7 iterations as shown in Fig. 5. Here, the algorithm found a consensus after iteration 5. In order to test the performance of the consensus, and navigation algorithms together, we carry out two tests. In the fist test, we simulated an environment with two targets (see Fig. 6b). Here, we did not introduce noise in the sensor readings (Fig. 6a. Around each target we placed a sub-swarm, the first one with three, and two agents respectively. The consensus results in both cases were positive (see Fig. 6c). Here, a positive identification is represented by 0 and a negative with X. In order to evaluate the robustness of system we carried out a second test, this time we added noise to the sensor readings (see Fig. 7a). The consensus in the swarm with less agents resulted negative, while the consensus in sub-swarm with more agents resulted positive (see Fig. 7b). This paper presents and consensus algorithm for the identification of targets in the context of swarm robotics. In particular, we present an application of target identification in which a swarm of drones is scanning a target area, looking for potential victims of a disaster, or targets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003146_022045-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003146_022045-Figure1-1.png", + "caption": "Figure 1. Texture configuration (a) the friction pair; (b) the shape and distribution and (c) the structure parameters of micro-texture", + "texts": [ + " The equivalent flow method is used to establish the average Reynolds equation which is solved by using the five-point differential and super-relaxation iterative method to calculate the coupling effect of different rough surface and bionic micro-texture on the lubrication performance of the cylinder surface. According to the working condition of the hydraulic cylinder, the gap between the cylinder and the piston is much smaller than the radius of the cylinder, just a few micrometers. So the cylinder can be unrolled into a plane for analysis and the diamond-like micro-texture is applied on the inner surface of the cylinder (Fig.1). To simplify calculations, a single control unit is analyzed and their structure parameters as follow in Table 1. When considering the random roughness on the surface of the friction pairs, the roughness on the inner surface of the cylinder and the outer surface of the piston is set to be \u03c31 and \u03c32 respectively. The gap between friction pairs is filled with the Newtonian oil. If the moving speed of the piston is U and the minimum gap thickness of the gap is h0, considering their direct contact, the roughness contact model is established as shown in Fig.2. IMMAEE 2018 IOP Conf. Series: Materials Science and Engineering452 (2018) 022045 IOP Publishing doi:10.1088/1757-899X/452/2/022045 Thus, the gap between the friction pairs is the thickness of the oil film in the flow field. When the roughness is not considered, it can be seen from Fig.1 that the thickness of the oil film can be expressed as 10 00 , , yxh yxhh h p (1) Where: \u21261 represents a non-diamond micro-textured region, and \u21260 represents a rhombic micro- textured region, consisting of 4 straight lines, namely: ab L ya L xbL ab L ya L xbL ab L ya L xbL ab L ya L xbL yx yx yx yx ) 2 () 2 (: ) 2 () 2 (: ) 2 () 2 (: ) 2 () 2 (: 4 3 2 1 (2) When considering the roughness, it can be seen from Fig. 2 that the actual oil film thickness becomes: 21 hhT (3) Assuming that the fluid can be characterized as Newtonian and slip phenomenon is not considered, the modified equation is given as follows by introducing the above parameters [9]: )(6)()( 33 xx h U y ph yx ph x s cyx (4) Where x and y are the pressure flow factors along the x and y directions, c is the surface contact factor and s is the shear flow factor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000443_detc2009-86926-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000443_detc2009-86926-Figure3-1.png", + "caption": "Figure 3. Hyper plane classifying two classes large margin", + "texts": [], + "surrounding_texts": [ + "Early prediction of faults in rotating machinery reduces maintenance costs as they are the major part of total operating costs in any industry. Condition-based maintenance and fault diagnosis have become important area of research that recommends maintenance decisions based on the information collected. The uses of condition based monitoring techniques have provided considerable savings in many industrial applications especially where large rotating machines are involved. Condition monitoring of rotating machinery helps in early detection of faults and anticipation of problems in time, so as to prevent complete failure. Recently, various artificial intelligence (AI) techniques such as hidden Markov models (HMM)[1], artificial neural networks (ANN) [2], and support vector machines (SVM) [3-4], have been used in the fault diagnosis of machines. Conventional pattern recognition method and ANN requires sufficient number of samples, which are sometimes difficult to obtain. SVM is based on structural risk minimization principle and has a very good generalization with few fault samples. So, it can produce high accuracy in fault classification. Samanta [5] has used adaptive neuro-fuzzy inference system (ANFIS) for prediction of features from vibration signals and compared it with support vector regression (SVR) and found that performance of (SVR) was better than (ANFIS) for the data sets used. Yuana and Chua [6] have used support vector machines for fault diagnosis of turbo pump rotor. They found that the \u2018one to others\u2019 multi-class SVM algorithm can correctly and effectively diagnoses multi-class faults in the turbo pump rotor test bed, and there is no reject region. In the present study, emphasis has been given towards condition monitoring of a rotor bearing systems by classifying and predicting various bearing faults using SVM, which is a statistical learning approach. Various bearing faults are simulated on machine fault simulator. Bearings vibration data are collected by piezoelectric accelerometers as time response. Features are extracted from time response and fed to SVM with the known output to train it. After training is done SVM can classify various faults based on given features. The signals obtained are processed for machine condition diagnosis as explained in the flow chart shown in Figure 1." + ] + }, + { + "image_filename": "designv11_92_0001079_icsma.2008.4505563-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001079_icsma.2008.4505563-Figure1-1.png", + "caption": "Fig 1. Mobile inverted pendulum", + "texts": [ + " As an extension of the previous research, in this paper, the boxing robot which has two arms to move as a mobile inverted pendulum structure is implemented. The robot has to keep balancing its own body as inverted pendulum systems and to navigate as the mobile robot. The structure of the boxing robot is quite different from the other mobile inverted pendulum systems in the literature[8- 11] because the robot is required to perform the boxing game. Experimental studies of controlling the mobile pendulum system for the boxing game are conducted. The wheel-driven mobile pendulum system is shown in Fig. 1. The pendulum is mounted on the mobile platform whose structure is a wheel-driven mobile robot. The mobile pendulum can change the direction by differing the wheel speed driven by DC motors. The control objective is to regulate the position tracking control of the cart while the pendulum is balancing. A gyro sensor is used to detect the angle of the pendulum and encoders are used to count rotations of wheels. Dynamic behaviors of the mobile pendulum are same as the inverted pendulum for a one dimensional case, but more uncertainties are present in the mobile pendulum system since rolling on the floor causes inconsistent disturbance to the system due to irregular surface of the floor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000348_12.828134-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000348_12.828134-Figure1-1.png", + "caption": "Fig. 1. Schematic of the device structures of (A) FEOET platform and (B) LOWE platform and droplet actuation principle.", + "texts": [ + " Since surface tension based droplet actuation provides a larger force than DEP, LOEW permits higher droplet transport speed in the order of cm/sec and allows more droplet manipulation functions such as droplet splitting or injection from a reservoir. Using LOEW, several important droplet manipulation functions have been demonstrated, including continuous 2D transportation, merging, mixing, and cutting. Both FEOET and LOEW have an open chamber configuration that allows easy integration with external tubing and reservoirs storing chemical and biochemical reagents. Fig. 1(A) shows the schematic of a FEOET device structure. It consists of a single-side glass substrate coated with two hydrogenated amorphous silicon (a-Si:H) layers, 0.5-\u03bcm undoped and 0.1-\u03bcm n+ doped, and a pair of 0.1-\u03bcm thick aluminum (Al) electrodes, separated by a 5-cm gap. An open poly-dimethylsiloxana (PDMS) chamber is fixed on top of the a-Si:H layer to house aqueous droplets and immiscible oil. A DC bias is applied to the Al electrodes, providing a lateral electric field across the entire FEOET device", + "aspx The device structure of LOEW is similar to FEOET, except LOEW uses a thinner dielectric layer between the photoconductive layer and the oil to allow more significant surface tension modulation through electrowetting. A thin 0.5-\u03bcm amorphous fluorocarbon polymer, Cytop (CTL-809M), layer is used in LOEW to increase the hydrophobicity and decrease contact angle hysteresis for droplet manipulation. During operation, a droplet immersed in oil is positioned on top of a LOEW chip with a voltage applied in the lateral direction. Under uniform light illumination, a droplet remains spherical (Fig. 1B(a)). When a shadow bar pattern is projected in the middle of the droplet, the contact angle at the two edges of a droplet decreases equally due to the enhanced voltage drop across the Cytop layer triggered by the shadow image (Fig. 1B(b)), but no net droplet movement is induced. However, when the image moves toward one side of the droplet, it breaks the originally balanced droplet shape and contact angle, resulting in a net droplet movement driven by the light-induced electrowetting effect. (Fig. 1B(c)). A simulation of the electric field distribution in the FEOET platform is performed using COMSOL Multiphysics 3.2. The device is simplified to two layers, a 5-\u00b5m a-Si:H, and a 350-\u00b5m electrically-insulating oil, to save computation time. A 100 V DC bias is applied at the two end planes to create a uniform and lateral electric field along the x-direction. The illumination of a 200-\u03bcm diameter circular laser beam is assumed to create a Gaussian photoconductivity distribution with a peak value 10-4 S/m" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000009_app.1962.070062201-Figure25-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000009_app.1962.070062201-Figure25-1.png", + "caption": "Fig. 25. (a) and T as function of time; ( b ) hysteresis loop; ( c ) complex modulus E, Young modulus El; loss modulus Ez, and 103s angle 6.", + "texts": [ + " Correlation between Stress and Strain Assuming that Both Units Vary Sinusoidally In many cases it can be assumed that the hysteresis loop resulting from the sinusoidal alternate strain on the cord sample tested is almost elliptically shaped. It follows that the alternate load shows an almost sinusoidal slope. The comparison of the planimetric area to that computed from the ellipse equation shows how far this holds true with the tests. On the basis of sinusoidal strain and stress conditions the following correlations are found according to Figure 25.lSJ7 Mi + (EZ/w)i + El\u20ac = ud sin wt (I) The general solution of eq. (1) is: e = ad sin (at - (o) / [ (E] - MU^)^ + E z ~ ] \u201d ~ (2) The loss factor tan 6 is, according to eq. (3) (Fig. 2%) : tan 6 = Ez/El (3) The angle \u2018p of the eq. (2) is given as: tan (o = E2/(El - Mu2) = tan 6/(1 - M d / E 1 ) (4) 2.3. Preload Elongation, Complex Modulus, Young\u2019s ModuZus, Viscosity Modulus, Specific Damping, and Loss Energy Formulae for the growth ev in the preload kv, for the complex modulus E, Young\u2019s modulus El, the viscosity modulus E2, the specific damping D, and the loss energy H (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002165_978-981-10-8672-4_18-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002165_978-981-10-8672-4_18-Figure1-1.png", + "caption": "Fig. 1 A schematic diagram of the mechanism with backlash", + "texts": [ + " First we consider the backlash model and its properties: Consider model of a mechanical system with backlash. This model combines with a robot manipulator dynamic model. Traditionally, backlash nonlinearity can be described by [15, 22, 33], q\u00f0t\u00de \u00bc p\u00f0t\u00de b if _p\u00f0t\u00de[ 0 and p\u00f0t\u00de q\u00f0t\u00de\u00fe b p\u00f0t\u00de\u00fe b if _p\u00f0t\u00de\\0 and p\u00f0t\u00de q\u00f0t\u00de b q\u00f0t \u00de otherwise 8< : \u00f01\u00de where q\u00f0t\u00de and p\u00f0t\u00de are output and input angles respectively, and _p\u00f0t\u00de is input angular velocity of a gearbox with backlash. The parameter b[ 0 is the backlash distance (Fig. 1). When backlash is presented in the joint gearbox, output angle follows input angle with a constant distance. If there is no contact in gears, output angle q\u00f0t\u00de will be between p\u00f0t\u00de b and p\u00f0t\u00de\u00fe b. In the proposed model the mentioned problem is avoided by the proposed adaptive robust controller. In almost all researches, input and output angles of the gears have been used in backlash modeling. However, in the proposed model, the input and output torques of the gear are used to simplify the overall model and the design procedure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001780_978-3-658-21194-3_110-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001780_978-3-658-21194-3_110-Figure8-1.png", + "caption": "Figure 8: Cutaway sketch of the Chaparral 2J, Falconer (2010).", + "texts": [ + " The large wing on the Chaparral 2E was an obvious attribute, however, to balance the downforce between the front and rear end another innovative idea was introduced, see figure 7. The front radiator opening was used entirely as an aerodynamic aid, and the front duct was carefully profiled to become a venturi to accelerate air passing through. The side panels provided endplates to optimize airflow over what was virtually a narrow wing span. In the late eighties this idea returned for efficient downforce in the front end of many sports car prototypes. In 1970 another ingenious vehicle from Chaparral was launched, the so called \u201csuckercar\u201d, Chaparral 2J, see figure 8. This car introduced the important concept of skirts for the sealing of the gap between vehicle and the road surface. The low pressure zone was enclosed underneath the car to press it towards the ground, just like a vacuum cleaner. To generate the low pressure zone an extra fan, driven by a two-stroke engine, was used. The 2J raced in the Can-Am series, it was quick, but unfortunately not very reliable and after complaints from fellow competitors the Chaparral 2J was banned. Though, the main contribution of this vehicle was the idea of enclosing the low-pressure zone using \u201cmovable\u201d sealing towards the ground" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000179_icbbe.2009.5162862-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000179_icbbe.2009.5162862-Figure3-1.png", + "caption": "Fig. 3 Schematic diagram of the FIA system for the determination of iron (II) or iron (III) a. carrier; b. PAR; c. potassium citrate; d. mixing coil; e. reaction coil; f. flow coil; D. detector; P1, P2. peristaltic pump; S. sample; V. sample valve; W. waste; PC. personal computer", + "texts": [ + " Dilute to 1l and mix thoroughly (1 ml=0.1 mg of Fe). More dilute solutions were prepared by diluting the stock solution. A model ZJ-la automatic metallic elements analyzer made by the Modern Analysis and Separate Technique Laboratory of Sichuan University and equipped with an optical detector and optical flow cell has the dual function of ion chromatography and flow injection analysis. The function of flow injection analysis was applied here to determine iron (II) and iron (III). The flow scheme is shown in Fig. 3. The flow system employed two peristaltic pumps (Shanghai Huxi Analytical Instrument Plant, P. R. China), which used to deliver all flow streams. The absorbance intensity was recorded at 718nm using an IBM-compatible computer. Date acquisition and treatment were performed with N-2000 software running under Windows XP. A series of working standard solutions with different concentrations between 1\u00d7 10-9 to 1\u00d7 10-5 g ml-1 were prepared by diluting a concentrated fresh standard solution of ferrous ammonium sulfate. As shown in Fig. 3, flow lines were inserted into sample or standard solution, potassium citrate, PAR, carrier solution, respectively. With the injection valve in the load position, the pumps were started to wash the whole flow system until a stable baseline was recorded. When the switching valve is in the washing position, the pump propelled working solution of iron (II) or samples to waste. Then the switching valve was rotated automatically to the sampling position so that iron (II) reacted with PAR. The absorbance signal was obtained by injecting 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001692_2018-01-1293-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001692_2018-01-1293-Figure4-1.png", + "caption": "FIGURE 4 Gear-shaft interference assembly resultant deformation- minimum interference (left), nominal interference (center), and maximum interference (right)", + "texts": [], + "surrounding_texts": [ + "slope deviation and transmission error has not been discussed previously in the literature.\nThe interference-fit (IF) joint typically consists of a gear body with teeth and rim assembled on a shaft and web and hub may or may not be present. The IF joint is an economical way of connecting the gear body to the shaft without the use of spline, bolts, or keys. It allows connection of a gear with a shaft of comparable diameter where there is not enough material to design a spline connection at low cost. Due to this inherent motivation for the IF joint, the gear rim thickness is an important design consideration to avoid thin rim effects such as- effect of rim thickness on root stress, effect of rim thickness on crack propagation, and the effect of rim thickness on mesh stiffness. The IF gear is usually designed to have a back-up ratio i.e. the tooth whole depth to the rim thickness ratio greater than 1.2. A typical gear design has the back-up ratio value between 1.2 and 2.\nFigure 1 shows an interference-fit gear on a shaft. Table\u00a01 shows the geometry details of the gear and the shaft for the reference design discussed in detail and henceforth called design A.\nThe gear-shaft joint in Figure 1 was analyzed by using Finite Element method.\nThe gear-shaft joint was analyzed using a Frictional contact type joint in ANSYS Workbench software. As the gear-shaft joint and the effect on gear are the interest areas of the analysis, only the portion of the shaft close to the gear mounting was included in the finite element model. The shaft was fixed at the centerline at one end and there was no other force acting on the model. The static structural analysis was carried out for the gear-shaft interference set at nominal, minimum, and maximum interference. X axis was oriented radially, Y axis denoted angular orientation, and Z axis was aligned along the centerline of the gear. FEA results are listed in Table 2.\nFigures 3 and 4 show the radial deformation along the X axis in the gear body for nominal interference. The deformation due to interference is maximum at near the interface and reduces slightly with increase in diameter. The maximum radial deformation increases with increase in interference.\nFigure 3 (center) shows the circumferential deformation in the gear body for nominal interference. The circumferential deformation due to interference is small and varies slightly from one end to the other. The maximum deformation increases with increase in interference.", + "\u00a9 2018 SAE International. All Rights Reserved.\nFigure 3 (right) shows the axial deformation in the gear body for nominal interference. The axial deformation due to interference is small. The maximum deformation shows slight increase with increase in interference.\nOverall, the gear body shows a significant radial increase in diameter with minimal changes in circumferential and axial directions.", + "The effects of the tooth distortion on the various aspects of the assembled gear are discussed further.\nISO 1328 [1] and ISO 10064 [2] give gear tooth profile and gear tooth helix measurement procedures. ISO 1328 defines the profile deviation as the amount by which the measured profile deviates from the design profile and defines the helix deviation as the amount by which the measured helix deviates from the design helix. The profile deviation is measured normal to the actual profile in the transverse plane in the profile evaluation range, whereas, the helix deviation is measured normal to the actual helix in the direction of the transverse base tangent in the helix evaluation range. Figure 5 shows the profile measurement details in relation with the gear geometry. The mean profile and the mean helix represent the shapes of the design profile and the design helix respectively and are aligned with the measured trace. Form deviation, slope deviation, and total deviation are used to characterize the measured trace.\n1\u00a0:\u00a0Design profile 2\u00a0:\u00a0Actual profile 3\u00a0:\u00a0Mean profile 1a\u00a0:\u00a0Design profile trace 2a\u00a0:\u00a0Actual profile trace 3a\u00a0:\u00a0Mean profile trace 4\u00a0:\u00a0Origin of involute 5\u00a0:\u00a0Tip\u00a0point 5\u00a0\u2212\u00a06\u00a0:\u00a0Usable profile\n5\u00a0\u2212\u00a07\u00a0:\u00a0Active profile C\u00a0\u2212\u00a0Q\u00a0:\u00a0Base tangent length to point\u00a0C \u03beC\u00a0:\u00a0Involute roll angle to point\u00a0C Q\u00a0:\u00a0Start of roll angle A\u00a0:\u00a0Tooth\u00a0tip\u00a0or start of chamfer C\u00a0:\u00a0Reference point E\u00a0:\u00a0Start of active profile F\u00a0:\u00a0Start of usable profile LAF\u00a0:\u00a0Usable length LAE\u00a0:\u00a0Active length L\u03b1\u00a0:\u00a0Evaluation range LE\u00a0:\u00a0Base tangent lenth to start of active profile F\u03b1\u00a0:\u00a0Total profile deviation ff\u03b1\u00a0:\u00a0Profile form deviation fh\u03b1\u00a0:\u00a0Profile slope deviation\nFigure 6 shows an involute tooth profile with a vertical shift after assembly overlaid on an involute tooth profile of an unassembled gear for comparison. The distance between the unassembled and the assembled profile as measured along the normal to the profile increases with increase in the roll angle and is a function of the radial deformation and the pressure angle at the point of measurement. The tooth profile curvature and the associated pressure angle change gradient with gear diameter, is greater near the root and decreases towards the tip with increase in the radius of curvature and the associated roll angle.\nd f dn r y= ( ),a (1)\ndn- Change in position along the normal to the involute. dr- Change in position along the radius. \u03b1y- Pressure angle at point y.\nThe resultant measured trace shows extra material on the profile that increases from the root to the tip resulting in positive profile slope. Table 3 shows the profile roll angle co-ordinates before assembly and the calculated estimate of profile deviation over the roll-angle range for normal radial deformation given by the FE analysis. As expected, the profile deviation due to radial distortion increases from the fillet to the tip. Figure 7 shows the expected profile measurements" + ] + }, + { + "image_filename": "designv11_92_0000489_imece2009-10196-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000489_imece2009-10196-Figure3-1.png", + "caption": "Figure 3. Example of a student-produced free body diagram of car undergoing event 4: tug of war.", + "texts": [ + " In order to illustrate some of the quality learning by the students, exhibits of post-lab deliverables (milestones) are presented in Figures 2-7. These examples of student works are presented chronologically according to their production throughout the semester. Considering that these are students at an early stage of their college education, the relatively high quality of the technical content of these exhibits (Figures 2- 7) is quite evident. Figure 2 is an example of a studentproduced project schedule performed on MS Project. Figure 3 is an example of a student-produced free body diagram of car undergoing event 4: tug of war (see below). Figure 4 is an example of a student-produced systems engineering Pugh matrix (for evaluating potential drive system designs for performance in different contest events). Figure 5 is example of a student-produced measured igure 6 is an example of a student-produced CNC machined measurement of the drive motor\u2019s characteristic (torque vs. speed) curve. F microcar chassis. 3 Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002206_s1068798x18070043-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002206_s1068798x18070043-Figure12-1.png", + "caption": "Fig. 12. Determining the chip-texture inclination \u03c8.", + "texts": [ + " The relative increase in Kc is calculated from the formula (22) The increase in Kc in the final cutting zone is accompanied by decrease in the inclination \u03a6 of the conditional shear plane, which is determined from \u2212\u0394 = c c.ex c.st c.st .K K K K RUSSIAN Eq. (6).2 Table 1 presents the results from Eq. (6) for \u03a6 in steady cutting and in the tool\u2019s exit zone. The relative decrease in \u03a6 is calculated from the formula (23) Experimental confirmation that Kc decreases is not possible by direct measurement, but indirect confirmation is possible by measuring the inclination of the chip texture and calculating the inclination \u03a6 of the conditional shear plane. The inclination \u03c8 of the chip texture (Fig. 12) depends on the inclination of the conditional shear plane (24) (25) Here \u03c7 is the angle between the tangent to the cutter side of the chip and the direction of chip texture (Fig. 12). We determine \u03a6 experimentally by means of a photograph of the chip texture in steel 20 on turning by a cutter with \u03b3 = 10\u00b0 (Fig. 13) [17]. In the calculations, we use Eqs. (24) and (25), with a mean angle \u03c7st = 59.65\u00b0 and with a value \u03c7ex = 76.16\u00b0 measured on a chip microsection (Fig. 13). The inclination of the conditional shear plane is found to be \u03a6st = 24.28\u00b0 in steady cutting and \u03a6ex = 13.2\u00b0 in the exit zone. From Eq. (5), when \u03a6st = 24.28\u00b0, \u03a6ex = 13.2\u00b0, and \u03b3 = 10\u00b0, we find that Kc.st = 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002415_gt2018-75782-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002415_gt2018-75782-Figure4-1.png", + "caption": "Figure 4. Computational grid and boundary conditions for ANSYS-CFX", + "texts": [ + " Within the current work only stepped-down labyrinth seals are considered corresponding to SH/RC>0 (Figure 3). The configuration of the stator is defined by the honeycomb height (HCH) and groove height (GH). Hereafter, HCH/RC=0 and HCH/RC>0 correspond to the smooth and honeycomb stator, whereas GH/RC>0 and GH/RC=0 to the stator with and without grooves above each fin tip respectively. It should be mentioned, that the center lines of groove and fin tip are axially aligned. The computational grid (Figure 4) consists of 1.0-2.0 million polyhedral cells depending on the configuration with a dense clustering towards the wall to ensure \u2206y+<1 in the entire domain, which allows integrating the governing equations down to the viscous sub-layer. 2 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/14/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use The computational model utilized within the general purpose CFD code ANSYS-CFX [11] employs the governing equation for the compressible gas with the properties depending on the local temperature, which are solved in the steady state mode with the k-\u03c9 RANS model [12] used for the turbulence closure. 3 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/14/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use The boundary conditions are specified as follows (Figure 4). The total temperature (tin), total pressure (pin), and corresponding turbulence quantities are specified at the inlet section. At the outlet section the static pressure (pout) is set, whereas the remaining transported quantities are extrapolated from the domain interior. At the rotor wall an adiabatic condition with the circumferential wall velocity (uwall) is employed, whereas a no-slip adiabatic condition is utilized at the stator and honeycombs walls (the closed honeycomb configuration is considered throughout the paper)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000700_cgiv.2009.70-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000700_cgiv.2009.70-Figure6-1.png", + "caption": "Figure 6. (a)-(c):full evaluation of a chessman model (314 patches) to depth 5 at 14fps; (d)-(f):full evaluation of the rocker arm model (354 patches) to depth 5 at 19fps; (a) and (d) are the original subdivision surfaces; (b), (c), (e) and (f) are deformed surfaces.", + "texts": [ + " By using the OpenGL Interoperability of CUDA, we can process these vertex buffers in GPU computation. After processing, it is directly rendered on GPU. The rendering process is highly accelerated by using vertex buffers. In order to measure the performance of our program, we wobble the vertices of the input mesh along their normals and reevaluate the mesh to the prescribed depth, i.e. 5 in our examples. Figure 1 shows reevaluating a toy model with 66 patches to depth 5 at 56fps. The ant model in Figure 7 with 298 patches is reevaluated to depth 5 at 14fps. More examples are shown in Figure 6(a) - Figure 6(c) on a chessman model with 314 patches and Figure 6(d)-Figure 6(f) on a rocker arm model with 354 patches. All these examples show that our patch-based tessellation algorithm achieves near realtime performance. Compared to our implementation on CPU, the GPU implementation runs about 20 times faster. For instance, the performance for the rocker arm model on CPU is less than 1fps. In this paper, a patch-based tessellation algorithm for DooSabin subdivision scheme is developed. Our patch-based tessellation creates a vertex patch for every vertex in the input mesh" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001904_012033-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001904_012033-Figure4-1.png", + "caption": "Figure 4. (a) Bending load and (b) Torsion load", + "texts": [ + " The parameters A, B, C, D, E, F and G [0, 1] are used to define shape of the bars, which can be either a round or a square tubing cross-sections as presented in Table 1, while the parameters H, I, and J are also used to define shape of bars, which can be either a round or a square bar cross-sections as presented Table 2. Moreover, a certain sections of the structure including H, I and J can determine a topology as displayed in Fig.3. Material density and young modulus of all bar are set to be 2x1011 N/ m2 and 7800 Kg/m3 respectively. The many-objective optimisation problem is posed to maximise natural frequencies of the first 5 modes, minimise structure mass and minimise construction cost subjected to stress and displacement constrains with 2 loading conditions, bending and torsion as shown in Fig.4. The optimisation problem can expressed as follow, Min: f1(x), f2(x), f3(x), f4(x) Subject to \u03c3bending max \u2264 \u03c3allow \u03c3torsion max \u2264 \u03c3allow -0.02 \u2264 y1 \u2264 0.06, m -0.025 \u2264 z1 \u2264 0.025, m -0.03 \u2264 y2 \u2264 0.06, m -0.06 \u2264 y3 \u2264 0, m -0.06 \u2264 y4 \u2264 0, m -0.06 \u2264 y5\u2264 0, m -0.06 \u2264 z2 \u2264 0.009, m 0.005 \u2264 t4 \u2264 0.02, m A, B, C, D, E, F, G, {0, 1} H, I, J {0, 1, 2} 3 1234567890\u2018\u2019\u201c\u201d where x is a vector of design variables. f1 and f2 are the maximum displacements in the y-direction (vertical direction) due to the bending load and the torsion load, respectively, while f3 and f4 are the inverse of the sum of natural frequencies in the first five modes and the structural mass, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001483_978-3-319-72730-1_17-Figure17.3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001483_978-3-319-72730-1_17-Figure17.3-1.png", + "caption": "Fig. 17.3 Field and current distributions in a rotating field machine", + "texts": [ + " In rotating field machines, the torque results from the interaction of a rotating sinusoidal field distribution b(x, t) and a rotating sinusoidal current layer a(x, t), as was explained in Chap. 3 of Part 1. As the speeds of both field distribution and 1For a brush axis in the neutral position, this requires a compensation winding or the absence of saturation-induced armature reaction; however, a brush axis which is not in the neutral position may result in armature reaction for non-rated operating conditions. current layer are the same, this torque is constant in time, but its magnitude depends on the angle \u03d1 between the symmetry axes of field and current layer (Fig. 17.3): T \u223c A\u0302 \u00b7 B\u0302 \u00b7 cos\u03d1 (17.3) For given current layer and field amplitudes, the torque is at a maximum when the symmetry axes of field and current layer are co-incident, i.e. \u03d1 = 0. When we draw the mmf distribution f (x, t) corresponding with the current layer a(x, t), it becomes clear that the mmf distribution due to the torque-producing current layer will not affect2 the field distribution if the symmetry axes of field and current layer are co-incident (\u03d1 = 0). We can therefore conclude that similar behaviour to that of the DC commutator machine can be obtained with a rotating field machine, on these conditions: 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003764_sdpc.2018.8664965-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003764_sdpc.2018.8664965-Figure1-1.png", + "caption": "Fig. 1. Rotor system model of tapered roller", + "texts": [ + " Some results and the conclusion are tested on the railway bearing comprehensive experimental platform. II. Dynamic model of tapered roller bearing rotor system A. System dynamics equation Taking the single wheel pair of the high speed train bogie as the research object, the wheelset includes the wheel shaft and the two wheels, and the eccentric mass of the rotor is equivalent to the center of the axle, and the bearing rotor system is formed. The dynamic model of a tapered roller bearing rotor system with eccentric mass is shown in Fig. 1. In the model, the left and right ends of the bearing rotor system are supported by the same bearing. The bearing is simplified as a combination of the spring system, and the dynamic model of the bearing is shown in Fig. 2. In figure 1, O1, O2, O3 respectively represent the bearing geometric center, rotor geometry center and rotor center of mass, the units are kg; K, C respectively represent the stiffness and damping of bearing, the units respectively are N/m and N\u00b7s/m; e represents the mass eccentricity of the rotor, the unit is m; FxL, FyL respectively represent the support counterforce in x direction and y direction of left end bearing, the units are N; FxR, FyR respectively represent the support counterforce in x direction and y direction of right end bearing, the units are N; In figure 2, m1, m2, mb respectively represent the quality of bearing inner ring, outer ring and unit resonator, the units are kg; x1\uff0cy1 respectively represent the lateral and vertical displacement of the inner ring of the bearing, the units are m; K1, C1 respectively represent the support stiffness and support damping of inner ring of bearing, the units respectively are N/m and N\u00b7s/m; x2 \uff0c y2 respectively represent the lateral and vertical displacement of the outer ring of the bearing, the units are m; K2, C2 respectively represent the support stiffness and support damping of outer ring of bearing, the units respectively are N/m and N\u00b7s/m; yb, Kb, Cb respectively represent the vertical displacement, stiffness and damping of the unit resonator, the units respectively are m, N/m and N\u00b7s/m; When bearing is damaged, periodic impact causes natural vibration of bearing inner and outer rings, sensors and other components" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000119_02602280910936183-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000119_02602280910936183-Figure5-1.png", + "caption": "Figure 5 A package-on-package formation, where two fully assembled and tested packages are stacked on top of each other. Each package may contain multiple stacked die", + "texts": [ + " D ow nl oa de d by O hi o U ni ve rs ity A t 1 8: 06 2 1 Ju ne 2 01 6 (P T ) memory, and die stacking makes the interconnections very short, improving the speed with which the processor can access the cache memory. Another common application for stacked die packages is USB memory sticks. Samsung makes a 16-chip memory stack package giving 16GB of memory. It uses 30mm thick wafers for the dies, and stacks 16 chips, each with 8GB of NAND logic. In the last few years, package-on-package stacking (PoP) has been introduced (Figure 5). Here, the packages themselves are designed so that they can be stacked on top of each other. Each package is fully assembled and tested. Provided the pad designs are compatible, different device types can be stacked. Toleno and Maslyk (2008) explains the drawbacks of die stacking inside a package, and the advantages of PoP. Within a package, the stacking is customised and fixed and requires a redesign of the entire die stack if there are any changes in any of the dies. If one die fails, the whole stack has to be scrapped" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003117_978-3-030-01382-0_6-Figure6.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003117_978-3-030-01382-0_6-Figure6.1-1.png", + "caption": "Fig. 6.1 Pendulum rotating in a vertical plane", + "texts": [ + " MATLAB Simulink simulation model in its most basic form consists of a single Simulink model that is self-contained: its input signals are generated internally using preexisting components library blocks, as well as the run commands, and the graphical outputs come from components such as Scopes that can be edited and manipulated entirely within Simulink. In the basic form, there is no need to use MATLAB script for controlling the Simulink model or for postprocessing of the simulation results. The basic ideas of how to set and run a Simulink model are explained in this section by means of examples of practical systems. We shall simulate an idealized friction-free model of the pendulum [2] shown in Fig. 6.1. The rod that has a length L is considered \u201cmass-less\u201d. The entire mass m of the pendulum is assumed to be concentrated in the bob. The pendulum rotates on a plane that is parallel to the page. The bob\u2019s rest position is h = 0. If we assign the bob an initial potential energy by setting it at an angle h0, it begins to swing back and forth. With no friction, the pendulum oscillates endlessly about h = 0. The amplitude of oscillations depends on the initial amount of potential energy. Another possible mode of operation happens whenever certain initial angle h0 combined with some initial angular velocity _h0 may cause the pendulum to tip over and rotate endlessly either in the clockwise or counterclockwise direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000501_fie.2009.5350615-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000501_fie.2009.5350615-Figure5-1.png", + "caption": "FIGURE 5 EXAMPLES OF SIMILAR APPROACHES TO WATER DISINFECTION BY AN ID STUDENT (LEFT) AND PET STUDENT (RIGHT), (COURTESY BERCH, BENSON)", + "texts": [ + " One challenge to this approach (each student doing their own design) was the substantial difference between the designs created by ID students and the designs created by PET students. The ID students\u2019 designs were much more refined in their functionality and usefulness by the intended user (children), but still had manufacturing challenges. The PET students\u2019 designs were simple to manufacture, but were very \u201cboxy\u201d, had little style, and may be difficult for a child to use safely. Some examples of these designs are in Figure 4 and Figure 5. Cascade Designs presented the project at \u201cDisinfection, 2009\u201d, a conference held in April, 2009 sponsored by the Water Environment Federation. In the paper published in the proceedings, the list of questions that Cascade Designs posed to the students during the final presentations is included. These questions included: \u2022 Does the form visually indicate how it is to be used, held and operated? \u2022 How stable is the design to prevent tipping? \u2022 Where are the fragile components (circuit board, solar panel, lights) located and are they adequately protected from breakage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002223_acc.2018.8431270-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002223_acc.2018.8431270-Figure1-1.png", + "caption": "Figure 1. Structure of the NHR.", + "texts": [], + "surrounding_texts": [ + "A. Simulation Results From ESO (8), positive constants \u03b5 and \u03c3 can affect the observation of unmeasurable states and disturbances deeply. In order to show the influence of constants \u03b5 and \u03c3, thenumerically simulated and ESO-estimated transient responses of some key process variables in the case of large-range power maneuver introduced above are given. This numerical simulation is done on Matlab/Simulink environment with a full-scale NHR model that much more complicated than model (1) utilized for ESO design. In this numerical simulation r1=100 and r2=300. Fig. 3 gives the simulated and estimated responses of normalized neutron flux nr, normalized delayed neutron precursor concentration Cri (i=1, 3, 5), average temperature of fuel elements Tf, average temperatures of primary coolant Tcav and secondary coolant Tlav, and reactivity induced by the control rods \u03c1r with \u03b5=0.1 and \u03c3=0.001, 0.1 and 0.02. Fig. 4 gives the responses with \u03b5=100, 500 and 1000 and \u03c3= 0.01. B. Discussions From Figs. 2 and 3, this reactor ESO can provide globally convergent bounded estimation for both the state-variables and disturbances under different values of parameters \u03b5 and \u03c3. From Fig. 2, it can be seen that for a given \u03b5, the variation of parameter \u03c3 mainly influences the estimation of external input reactivity \u03c1r and thermal-hydraulic variables Tf, Tcav and Tlav. Furthermore, \u03c3 is smaller, the transition time of the estimated states to the bounded set cover the simulated states are smaller. From Fig. 3, for a given \u03c3, the variation of parameter \u03b5 mainly influences the estimation of external input reactivity 3000 4000 5000 6000 7000 Time/s 3000 4000 5000 6000 7000 Time/s 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 n r 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 C r1 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 C r3 Time/s 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 C r5 Time/s Simulation = 100 = 500 = 1000 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 n r 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 C r1 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 C r3 Time/s 3000 4000 5000 6000 7000 0.2 0.4 0.6 0.8 1 C r5 Time/s Simulation = 100 = 500 = 1000 3000 4000 5000 6000 7000 200 300 400 500 T f/\u2103 3000 4000 5000 6000 7000 200 220 240 260 T c a v /\u2103 3000 4000 5000 6000 7000 -3 -2 -1 0 \u03c1 r/\ufe69 Time/s 3000 4000 5000 6000 7000 200 210 220 230 T la v /\u2103 Time/s Figure 3. Numerical simulation results with \u03c3=0.01 and different \u03b5. \u03c1r and neutron-kinetic state variables nr and Cri (i=1, \u2026, 6), and the estimated states converge to the bounded set around the simulated states faster if \u03b5 is smaller. Actually, from inequality (24), if either positive constants \u03b5 or \u03c3 is smaller, the total norm of the negative parts in the right side of this inequality is larger, which leads to a faster convergence. Moreover, from the responses of state-variables especially those of disturbances \u03c1r and Tlav, the steady estimation errors in a lower power-level are larger than those in a higher power-level, which can also be interpreted from inequality (24). Actually, the reactor power-level is lower, normalized neutron flux x1=nr is smaller, which further induces a larger norm of the positive terms in the right side. Then, it can be seen from (25), the positive terms are stronger, the set is larger, which finally leads to a higher disturbance estimation error. Finally, it can be seen from the above numerical simulation results and discussions that the simulated responses of state-variables and disturbances are well in accordance with the theoretical result, which verify the correctness of the ESO design for the widely-used nuclear fission reactors." + ] + }, + { + "image_filename": "designv11_92_0000738_acc.2009.5160109-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000738_acc.2009.5160109-Figure1-1.png", + "caption": "Figure 1 - Combustion Engine Test Bench System", + "texts": [ + " The Section III explains the structure of the observers used (EKF, HGO, SMO and NESO). In the Section IV simulation results of these observers in open-loop are presented. A comparison in terms of mean quadratic error is also performed. A comparison in closed-loop, between two of these observers and the partial observer of [1], is presented in Section V. Finally, conclusions are given in Section VI. A 978-1-4244-4524-0/09/$25.00 \u00a92009 AACC 4648 The typical structure of the combustion engine test bench is illustrated in Figure 1. \u03b1 The main parts of such a system are the dynamometer, used to simulate the load, the connection shaft and the combustion engine itself. The engine and dynamometer speed are measured using incremental encoders and, therefore, are known quantities. The engine torque and the torsion angle are hard to measure and therefore unknown. A typical control design objective for such a system is the reference tracking of the engine torque and speed, by controlling the air gap torque of the dynamometer and the throttle pedal of the combustion engine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.2-1.png", + "caption": "Fig. 11.2 Comparison of theoretical tire models for cornering. Reproduced from Ref. [2] with the permission of Guranpuri-Shuppan", + "texts": [ + " Performance criteria of the force and moment of tires are introduced to improve the cornering performance of a vehicle. Furthermore, the effect of alignment on the cornering performance of a vehicle is explained by the force and moment of tires. Tire models of vehicle dynamics are categorized as theoretical and empirical models. Theoretical models are mechanical models used to calculate the force and moment of a tire, and they can be further categorized as several types of model as shown in Fig. 11.1. The solid model in Fig. 11.2a does not consider deformation of the carcass or tread ring (belt), the rigid ring model in Fig. 11.2b considers deformation of the carcass but not of the tread ring, the flexural (elastic) ring model in Fig. 11.2c considers deformation of the carcass and tread ring, and the finite element (FE) model in Fig. 11.2d is used in FEA [1, 2]. The treads of all theoretical models except the finite element model are modeled using the brush model in which \u00a9 Springer Nature Singapore Pte Ltd. 2019 Y. Nakajima, Advanced Tire Mechanics, https://doi.org/10.1007/978-981-13-5799-2_11 707 small elements of tread rubber are deformed independently of the direction of the applied force like bristles of a brush. The cross-sectional deformation of the solid model is shown in the right figure of Fig. 11.2a. The side force is generated by shear deformation of the tread rubber, and the shear displacement increases until the sliding point at which the shear force is equal to the static frictional force. The tread slides laterally after passing the sliding point, and the tread moves back to the original position at the trailing edge. The side force of a tire is expressed by Side force \u00bc Z A lateral spring rate of tread per unit area\u00f0 \u00de \u00f0displacement of tread\u00de dA; \u00f011:1\u00de where A is the contact area. In the rigid ring model of Fig. 11.2b, the tread ring behaves like a rigid body. However, the ring can deform in torsional and lateral directions because the ring is supported by carcass springs. When the rigid ring model rolls with the slip angle a, there is a lateral displacement and lateral slip. A comparison of Fig. 11.2b with Fig. 11.2a shows that the differences are the rigid body displacement y0 of the center of the ring and torsional displacement. The side force of the rigid ring model is also given by Eq. (11.1). The adhesion region is from point O to point A, while the slip region is from point A to the trailing edge (x = l). Meanwhile, the string model is used in the transient analysis of a vehicle for simplicity. The elastic ring model of Fig. 11.2c is used for the analysis of a radial tire. When the belt is represented by a beam, the model is called a beam model. Fiala [3] approximated the in-plane deformation of the beam using a parabola and neglected the effect of belt tension; his model is called the Fiala model. When the elastic ring model rolls with slip angle a, the tread shear deformation of the elastic ring model shown by slanted lines is smaller than that of the rigid ring model owing to the bending deformation of the tread ring. Hence, the side force of the elastic ring model is weaker than that of the rigid ring model", + " Various physical tire models have been proposed. SWIFT was developed by TNO on the basis of the studies of Zegelaar [17] and Maurice [18], F-Tire was developed by Gipser [19\u201322], RMOD-K was developed by Oertel and Fandre [23], and the Hankook tire model was developed by Gim et al. [24\u201327]. Those models are used for describing not only cornering phenomena but also the riding comfort of the vehicle/suspension/tire system. Lugner et al. [28] published a paper comparing these models. The finite element model as shown in Fig. 11.2d presents the tire shape and construction in detail and can be used to quantitatively predict the cornering performance of a tire. However, because the finite element model is computationally expensive, the application for the vehicle dynamics is limited and the model is mainly applied in tire development. Meanwhile, an empirical model is a mathematical model that fits measurements made in an indoor test and is used for the vehicle dynamics. The most famous model is the Magic Formula proposed by Pacejka [1]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure7-1.png", + "caption": "Fig. 7 Calculating displacement during acceleration", + "texts": [ + " Another advantage of the proposed method emerges when a deflection analysis is performed. The deformation direction may be observed in the wick coupling stiffness direction instead of the real deformation direction because of the coupling effect of 1D elements (for spring elements) and 3D elements (for the shaft). Lastly, explaining the simulation results using the spring element method is difficult, whereas using the newly proposed method solves the usual problems related to the DOF coupling effect. Figure 7 shows the results of using the two methods to calculate the rotor-bearing system\u2019s displacement and deformation during acceleration. A comparison of the results obtained by Methods 1 and 2 is shown in Figs. 8(a) and 8(b), respectively. When the spring element-based method is used, the maximum displacement is in the circumferential direction (Fig. 8(a)), which is obviously not consistent with reality. However, when the proposed plate theory-based method is used under the same load conditions, the direction of maximum displacement (0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003115_ecce.2018.8557727-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003115_ecce.2018.8557727-Figure11-1.png", + "caption": "Fig. 11. 3D thermal model under fault condition.", + "texts": [ + "5 pu while the RMS phase currents in terminally shortcircuited ABC set are also quite low, being 28.8A, 45.8A and 47.9A, respectively. Consequently, total heating effect in the ABC set is lower than that of the other two healthy sets. It has been calculated that the copper loss of each healthy 3-phase set is 3.5 times larger than that of the faulty 3-phase set. Because the asymmetric loss distribution, it is essential to use the full 3D thermal model containing 36 slots and one half of the axial length as shown in Fig. 11 for accurate thermal analysis. The temperature distribution under one turn SC with 3-phase terminal SC when the machine operates at 4000rpm with 120A current in healthy phases is shown in Fig. 12. It is shown that the two healthy sets have nearly symmetric temperature distribution and their overall temperature is higher than that of the faulty set. The hotspot is located in the middle part of the end windings of the healthy 3-phase sets, similar to those seen in Figs. 8 and 9. Table III compares the steady state temperatures obtained by the predictions and measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002338_s11538-018-0460-0-Figure17-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002338_s11538-018-0460-0-Figure17-1.png", + "caption": "Fig. 17 (Color figure online) Graphical representation of the F\u00b1 that map the domain of dependence B0 (41) to the lamellipodium. The s = 0 boundary of B0 is mapped to the membrane of the cell and the s = \u2212 1 to the minus ends of the filaments inside the cell. The filaments and the other functions of \u03b1 are periodic with respect to \u03b1. The \u201cfilaments\u201d plotted in the lamellipodium correspond to the interfaces of the discretization cells of B0 along the \u03b1 direction", + "texts": [ + " We numerically solve the FBLM (6) with a problem-specific FEM that was first presented in Manhart et al. (2016). Here we present some of its components. The maximal filament length varies around the lamellipodium, and in effect the computational domain B(t) = {(\u03b1, s) : 0 \u2264 \u03b1 < 2\u03c0, \u2212L(\u03b1, t) \u2264 s < 0} is non-rectangular. For consistency and stability reasons we recover the orthogonality of the domain B(t), using the coordinate transformation (\u03b1, s, t) \u2192 (\u03b1, L(\u03b1, t)s, t) , and replace it by B0 := [0, 2\u03c0) \u00d7 [\u2212 1, 0) (\u03b1, s). (41) See also Fig. 17. Accordingly, the weak formulation of (6), recasts into 0 = \u222b B0 \u03b7 ( \u03bcB\u22022s F \u00b7 \u22022sG + L4\u03bcA D\u0303tF \u00b7 G + L2\u03bbinext\u2202sF \u00b7 \u2202sG ) d(\u03b1, s) + \u222b B0 \u03b7\u03b7\u2217 (L4\u03bc\u0302S ( D\u0303tF \u2212 D\u0303\u2217 t F \u2217) \u00b7 G \u2213 L2\u03bc\u0302T (\u03c6 \u2212 \u03c60)\u2202sF\u22a5 \u00b7 \u2202sG ) d(\u03b1, s) \u2212 \u222b B0 p( ) ( L3\u2202\u03b1F\u22a5 \u00b7 \u2202sG \u2212 1 L \u2202sF\u22a5 \u00b7 \u2202\u03b1(L4G) ) d(\u03b1, s) + \u222b 2\u03c0 0 \u03b7 ( L2 ftan\u2202sF + L3 finnV ) \u00b7 G \u2223\u2223\u2223 s=\u22121 d\u03b1 \u2213 \u222b 2\u03c0 0 L3\u03bbtether\u03bd \u00b7 G \u2223\u2223\u2223 s=0 d\u03b1, (42) with F,G \u2208 H1 \u03b1 ( (0, 2\u03c0); H2 s (\u2212 1, 0) ) . In a similar manner the modified material derivative and inextensibility conditions read D\u0303t = \u2202t \u2212 ( v L + s\u2202t L L ) \u2202s and |\u2202sF(\u03b1, s, t)| = L(\u03b1, t)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002845_ipec.2018.8507656-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002845_ipec.2018.8507656-Figure1-1.png", + "caption": "Fig. 1. Photograph and cross-sectional view of test motor", + "texts": [ + " This current profiling method worked well at light load condition. However, it did not work well at high load condition over rated load especially at current overlapping region. In this paper, a reason why this method in [8] (here after, conventional method) does not work is investigated through the magnetic analyses using 2D-FEA. Then a novel current profiling method considering mutual coupling effect is proposed. Finally, through the simulation and experiment of a 400W 4-phase 8/6 SRM that has shape of Fig.1, it is proven that the torque ripple minimization by using the proposed method is working well at wide load condition. The process of torque ripple minimization is explained. Firstly, by using torque contour function, the target phase torque is generated. Then, through the current-torqueposition (here after, i-T-\u03b8) function, the target phase current profile required is computed. Finally, through PWM voltage control, the target current profile is realized. As can be seen Fig. 2, the torque contour function selected among many kinds of possible functions [2] as an 3418 \u00a92018 IEEJ example is explained for 4-phase SRM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002422_aim.2018.8452315-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002422_aim.2018.8452315-Figure1-1.png", + "caption": "Figure 1. 7axis manipulator system(left) and each axis (right)", + "texts": [ + " 2 3 2 1 2 1 1 1 2 1, 2 1, \u02c6 [ / ( / ) ( / ) \u02c6\u02c6] ( / ) ( ) j j j j oj j j oj j oj j j oj j dj j j dj j s u k c k k x k x x c x x The control input \ud835\udc62?\u0305? can be expressed as Eq. (12). 2 2 1 2 1 1 3 1 2 1, 2 1, 1 {[ / ( / ) ( / ) ] ( / ) ( ) \u02c6\u02c6( ) } j j oj j j oj j oj j j j oj j j j dj j j dj j u k c k k x k x K sat s x c x x The chattering can be reduced by compensating the estimated perturbation value as shown in the above equation (12) to the control input. In this paper, we use 7-axis manipulator system. System hardware structure and position and direction of rotation for each axis are shown in Figure 1. The system configuration diagram of the 7-axis manipulator is shown in Figure 2, and 978-1-5386-1854-7/18/$31.00 \u00a92018 IEEE 34 the main computer will output the digital control input and the analog input / output board will convert the digital control input to the 10V analog control input. The analog control input is applied to the BLDC motor through the motor driver to control the robot manipulator. B. Inverse Kinematics In order to move the end effector of the manipulator to the desired position, the inverse kinematics, which is the method of finding the angle of each motor relative to the position of the end point, must be solved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000100_cp:20080617-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000100_cp:20080617-Figure2-1.png", + "caption": "Figure 2: Description of 'Sensor angle'. (a) Instant of switching on of 'Tl' for s =0\u00b0. (b) Instant of switching on of 'Tl' for an arbitrary positive value ofs.", + "texts": [ + " The rotor position is defined here via the angle variable Or, which is the space angle between the axis of the 'A' phase winding of the PMSM's armature and the rotor's 'q' (quadrature) axis. Positive rotor rotation is assumed in anticlockwise direction and the rotor's 'q' axis is assumed to lead the field (i.e. 'd' or direct) axis. If Figure 4: Simulated steady state wave forms of a particular phase current and DC link current at, Vdc = 35 Volts,j= 0.0016Nm-second/Radian, TL=1.5 Nm, Sensor anele. S = 0\u00b0. proper setting of the position sensor and the 'Inverter Controller' ensure that transistor T1 of Fig. 1 becomes on when Or = -60\u00b0 (Fig. 2(a)), current will start entering armature 'A' phase. Under idealized assumptions, current is assumed to build up instantaneously in that phase. From this position, after 120\u00b0, T1 will become off, when Or = +60\u00b0 and current in 'A' phase is ideally assumed to become zero instantaneously. Such a case of synchronization between inverter device switching and rotor space position is defined as the reference sensor position for which the sensor angle value, s, is defined zero. If, by proper adjustment of the position sensors and the Inverter Controller, the transistor T1 is made on when Or = ( 60+s)0, as shown in Fig. 2(b), the sensor angle for such a situation is in a lagging sense and is equal to +so. For a leading sensor position, value of s will be negative. For the sensor position of s = 0\u00b0, therefore, the fundamental component of any armature phase current and the fundamental component of the corresponding phase's induced EMF will be co-phasal. Hence the fundamental armature MMF phasor and the field MMF phasor will be orthogonal, resembling a conventional DC machine. 4 The Detailed Numerical Model The 'Detailed numerical model' (DNM), developed here, is based on iterative method of solving the PMSM's non-linear differential equations [2], taking into consideration, the cyclically changing patterns of the conducting inverter power devices as per 120\u00b0 conduction mode of the inverter", + " But, in the implemented PMSM laboratory prototype, the back emf profile is experimentally found to be flat topped with the flatness span of about 540 electrical. This can be proved from the waveform of Fig. 7. It shows the steady state experimental waveforms of the induced EMF ofand is found to be linear. Now, it can be concluded that, Or = mrt + OrO (7) Where, Oro == Value of electrical rotor position, Or when ia attains its peak value. If the concept of sensor is taken into account as shown in Fig.2, it can be concluded that:\u00b0ro = S (8) Now, using (8) in (6), term3 of equation (6) becomes: Term3= 1.655lj10 iUnk COS(S)mr (9) With the help of (6) and (9), (5) takes the ultimate form: -2 Po; =1.8~ilink +1.82Lsilin,pilink+1.65i;/oilinkcos~)UJ,. (10) Now, the voltage at the front end, DC link voltage of the inverter is, vde. Using assumption (v), it is obtained: P0; = Vde i link (11 ) Using (10) and (11), it can be written: - - Vde = 1.8~4ink+1.8~pilink+1.65ijtoCOS~m (12) This equation looks similar as that of the voltage equation of the armature circuit of a conventional DC motor with constant excitation, where iUnk may be thought of as the equivalent DC machine's armature current, flowing through the equivalent DC machine's armature of resistance value equal to 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002302_saci.2018.8440978-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002302_saci.2018.8440978-Figure2-1.png", + "caption": "Fig. 2. Rigid body process model.", + "texts": [ + " the electromechanical process which contains: a brushless direct current (DC) drive motor, a brushless DC disturbance motor, a drive disk, a load disk, a speed reduction assembly, high resolution encoders at drive and load discs, elastic or inelastic drive belts. The two disks have variable moment of inertia which may be adjusted by adding or removing brass weights. The laboratory equipment is used for the validation of the designed control solutions. The process is modelled with rigid body dynamics as in Fig.2. The dynamics that describe the process can be written as: .)(] ' )([ ,)(])()([ . 22 2 1 .. 2 2 2 . 1 2 21 .. 1 22' Dldpdr Dldpdr grTcgrcJ gr grJgrJ TgrccgrJgrJJ =++++ =++++ \u2212\u2212\u2212 \u03b8\u03b8 \u03b8\u03b8 (1) with Jdr, Jld, Jp, gr and gr\u2019 expressed as: .', 6 , , _ _ _ _ _ __ __ __ dr dp lp dp drlp dpld backlashldpdrpp ldwlddld drwdrddr n n gr n n nn nn gr JJJJ JJJ ,JJJ === ++= += += (2) where: Jdr, Jld, Jp \u2013 the inertia of the drive disk, load disk and pulley inertia, c1, c2 \u2013 the drive and load friction, gr \u2013 the drive gear ratio, gr\u2019 \u2013 the partial gear ratio between the drive disk and the idler pulley, 1, 2, p \u2013 the drive/load disk and idler pulleys positions where 1=gr 2 or 1=gr\u2019 p Considering only 1 as the process output, the states can be expressed as: x1= 1 and x2=d 1/dt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003764_sdpc.2018.8664965-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003764_sdpc.2018.8664965-Figure4-1.png", + "caption": "Fig. 4. Outer ring fault model", + "texts": [ + " According the equation above, the total nonlinear contact force of the bearing in the x- and y-directions are 100 9 xR 100 9 yR cos 1 sin 1 t i i i t i i i N F K H i N F K H i (9) Supposing the outer bearing ring of the right-hand bearing is faulty. \u03b8f represents the local of the defect center. \u03b8cr, located in the bearing area, represents the center angle corresponding to the arc of the defect. L0 represents the fault wide. When the roller passes through the fault location, the normal contact deformation of roller at the angular displacement \u03b8f and the raceway will change, the change is couter. rg represents the radius of the cross section of the centroid of a tapered roller, as is shown in the Fig. 4, the conditions of the change in a period of time is \u03b8f \u2013 \u03b8cr / 2 < mod(\u03b8i , 2\u03c0) < \u03b8f + \u03b8cr / 2, that is range condition. couter is defined as 2 2 0 g g outer range condition 2 0 L r r c else \uff0c \uff0c (10) The variation of deformation of tapered roller in the fault area will lead to the change of nonlinear Hertz contact force. Because of the change, the amount of the deformation is r1 r2 r1 r2 0 outer( )cos ( )sini i ix x y y c c (11) D. System model parameters The main parameters of rotor system and unit resonator selected in this paper are shown in table I and table II" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001780_978-3-658-21194-3_110-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001780_978-3-658-21194-3_110-Figure12-1.png", + "caption": "Figure 12: Geometrical definition of the \u201cflat\u201d underbody.", + "texts": [ + " If the contact between the vehicle and road surface was lost, the downforce generating pressure was equalized with the speed of sound so the downforce was lost immediately and the car was impossible to manoeuvre. After a number of serious accidents, for instance the fatal crash of Gilles Villeneuve at Zolder in Belgium 1982, the ground effect technology was banned in 1983. A new regulation was introduced that mandated formula one cars to have a flat bottom between the inside tangent of the tyres, as is shown in figure 12. No skirts were allowed. The effect of this change in regulation was massive, and it would mean a revolution in the downforce design. At first, there was a reduction of the cornering speed, however, simultaneously this would be the \u201ckick off\u201d for an intensification of the development for increased downforce through manipulation of the under body. From a fluid mechanics perspective, devices for downforce generation had hitherto been relatively simple: add on wings at different locations, curvatures of selected surfaces and mechanical arrangements to eliminate pressure equalization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002733_978-94-007-6046-2_18-Figure41-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002733_978-94-007-6046-2_18-Figure41-1.png", + "caption": "Fig. 41 FEM analysis", + "texts": [ + " The joint torque sensors using the CrN-STFs have the following distinguished characteristics: \u2022 High sensitive and high stiff torque sensing \u2022 Precise sensing by precise locating of tiny sensor elements \u2022 Reliable torque sensing by redundant strain bridge circuits \u2022 Stable and high-resolution torque sensing by integrated sensing elements as described in section B We have developed four different types of the joint torque sensors in different capacity of sensing torque. They are designed to be implemented in the torqueservo modules. Table 3 shows their mechanical properties. The measuring range is a double load of the rated torque of each torque sensor. Figure 41 shows a sample FEM analysis of the Type B torque sensor. Stress-strain behavior is simulated under the condition of rated torque load on the inner race with fixed outer race. Displacement between the fastener point of the inner race and one of the outer race is less than 0.0025[mm] under the rated torque load of 30[Nm]. Torsional stiffness of the Type B torque sensor is 3:04 104[Nm/rad]. The redundant strain bridge circuits are dual, group A (R1-R2-R7-R8) and group B (R3-R4-R5-R6) in four active gauge methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001730_icomet.2018.8346421-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001730_icomet.2018.8346421-Figure1-1.png", + "caption": "Fig. 1. Helicopter free body diagram = + (1) = + (2) = + (3)", + "texts": [ + " The non-linear model is obtained with modest level of complexity using first principle approach. The different controllers designed in this work include Increment-Proportional-Integral-Derivative (IPID) controller and H mixed sensitivity design. These controllers effective design approach can be judged from the simulation results. II. ATTITUDE DYNAMICS The derivation of transfer function for attitude dynamics of a small scale helicopter is carried out during its flight at low speed and hovering. The free body diagram of small scale helicopter is shown in fig. 1. OXYZ describes body frame and represents inertial frame. Main rotor and tail rotor thrust vector are described by and respectively. The anti-torque of main rotor is constant during hover and , similarly for tail rotor. The tilt-angles for longitudinal and lateral dynamics are described with respect to rotor hub plane by \u2018a\u2019 and \u2018b\u2019. Equations describing attitude dynamics are given as [18]; Where , and represents fuselage inertia about x-, yand z-axis respectively; \u03c9 = describes angular velocity components and L, M and N denotes the roll, pitch and yaw moments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000764_eej.20631-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000764_eej.20631-Figure2-1.png", + "caption": "Fig. 2. Magnetic rail.", + "texts": [ + " The poles of the permanent magnets are the same as those of the magnetic rail, and a repulsive force is generated. When the load becomes large, the weight stage descends. The distance between the permanent magnet on the weight stage and that on the rail becomes small, and the weight of the load and the repulsive force come into balance. Thus, a large levitation force is obtained from the repulsive force of the permanent magnet and we can vary the weight without changing the levitation gap of the carrier body. Figure 2 shows the magnetic rail and Fig. 3 shows the permanent magnet for the carrier. Table 1 shows the dimensions of the permanent magnet. A neodymium magnet is used. Table 2 shows the dimensions of the HTSC. YBaCuO material is used. First, the shape of the HTSC is considered in order to set the dimensions of the system. The levitation gap range is defined. The dimensions of the magnetic rail, such as the pole pitch and the width of the magnet, are calculated so that flux of the permanent magnet is pinned efficiently by the HTSC. The magnetic rail shown in Fig. 2 achieves the most efficient pinning effect at a levitation gap g0 = 3 mm. 2.2 Experimental method The HTSC was indirectly cooled for 10 minutes and was placed on iron gauze. Liquid nitrogen was placed under the iron gauze. The cooled air from the liquid nitrogen cooled the HTSC indirectly. Then the HTSC was set on the carrier. An acrylic spacer was set on the magnetic rail, and the carrier was placed on the spacer. Liquid nitrogen was poured into a vessel for the HTSC, and the HTSC was cooled directly", + " When a rolling angle occurs, the levitation gap is also changed and the stability of the pinning force is disturbed. In the previous experiment, extra weight was added under the carrier to increase the inertia of rolling motion. The center of gravity of the carrier was also lowered to increase the stability to rolling motion. Next, a more stable configuration of the system with two magnetic rails and four HTSCs is considered. Figure 8 shows the configuration and the appearance of the four-HTSC support-type system. Two magnetic rails are used, as shown in Fig. 2, and the number of the HTSC is doubled. The load characteristic of the four-HTSC configuration is shown in Fig. 9. The weight from 30.6 N, which is the weight of the weight stage, is measured. The weight stage for the four-HTSC configuration is slightly heavier than double that of the two-HTSC configuration. When the number of weight stages is increased to two, the displacement for the same weight becomes half that of the two-HTSC configuration. Although Fig. 9 shows measurement results up to 80 N, no rolling motion was observed even at 150 N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.12-1.png", + "caption": "Fig. 11.12 Top view of the string and equilibrium of forces in the contact region [29]", + "texts": [ + "11 Force equilibrium in a tread element of the string model [29] where v is the displacement of the string in the y-direction and ks, S1, D and fy are, respectively, the lateral carcass spring rate per unit length in the circumferential direction, the tension of the string in the circumferential direction, the shear force and the side force on the contact patch. The shear force D is proportional to the shear angle \u2202v/\u2202x: D \u00bc S2 @v=@x; \u00f011:24\u00de where S2 is the tension related to the shear deformation. We here introduce the equivalent total tension S (=S1 + S2). Using Eqs. (11.24) and (11.23) can be simplified as S @2v=@x2 ksv \u00bc qy: \u00f011:25\u00de Because fy is zero outside the contact patch, Eq. (11.25) can be rewritten as S @2v=@x2 ksv \u00bc 0 for xj j[ a; \u00f011:26\u00de where a is half the contact length. Figure 11.12 shows the deformation of the string model near the contact patch. We also introduce the relaxation length r expressed by r \u00bc ffiffiffiffiffiffiffiffiffi S=ks p : \u00f011:27\u00de Using Eqs. (11.27) and (11.26) is rewritten as r2 @2v=@x2 v \u00bc 0 for xj j[ a: \u00f011:28\u00de If the relaxation length is shorter than the tire circumference, the displacement at the trailing edge v2 is not affected by the displacement at the leading edge v1. Displacements of the string outside the contact patch can be independently solved at the leading and trailing edges: v \u00bc C1e x=r for x[ a v \u00bc C2ex=r for x\\ a: \u00f011:29\u00de Considering the boundary conditions v = v1 at x = a and v = v2 at x = \u2212 a, we obtain @v=@x \u00bc v1=r for x # a @v=@x \u00bc v2=r for x \" a; \u00f011:30\u00de where x#a denotes that x gradually decreases toward a while x\" \u2212 a denotes that x gradually increases toward \u2212a. Considering that the slope of the string is continuous between the inside and outside of the contact patch at the leading edge, we obtain @v=@x \u00bc v1=r for x \" a; x \u00bc a; x # a: \u00f011:31\u00de (2) Force and moment of the string model The force and moment of the string can be calculated from the displacement in the contact patch v, the lateral carcass spring rate per unit length ks in the circumferential direction and the tension of the string S as shown in Fig. 11.12. The side force Fy is given by Fy \u00bc ks Za a vdx\u00fe S v1 \u00fe v2\u00f0 \u00de=r: \u00f011:32\u00de The self-aligning torque around point O is obtained through the addition of the lateral and circumferential components. The self-aligning torque due to the lateral displacement of the string v is given by M0 z \u00bc ks Za a vxdx\u00fe S a\u00fe r\u00f0 \u00de v1 v2\u00f0 \u00de=r: \u00f011:33\u00de The self-aligning torque due to the circumferential displacement u is given by M z \u00bc Cx Za a Zb 2 b 2 uydxdy; \u00f011:34\u00de where Cx is the shear spring rate of the tread per unit area in the circumferential direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002435_978-3-319-99522-9_3-Figure3.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002435_978-3-319-99522-9_3-Figure3.1-1.png", + "caption": "Fig. 3.1 Composed motion of a point", + "texts": [ + " The motion of the latter moving frame can either be predefined as a target motion or is obtained directly as a particular solution from the equations of motion. The position, the velocity and the acceleration of a point depend on the reference coordinate system, relative to which the motion of the point is analysed. The motion of a point will be described differently by two observers which are moving relative one to another. We consider a fixed frame of reference X ng1 and a second frame Oxyz which move in a known, but arbitrary, fashion with respect to a fixed base (Fig. 3.1). Let P be a particle moving in both spaces of the two frames, as its position is defined at any instant by the vector ~q in the fixed frame and by the vector ~r in the moving frame. By denoting~qO the position vector of the origin O of the moving frame with respect to the fixed frame, we obviously have ~q \u00bc~qO \u00fe aT~r; \u00f03:1\u00de where a is the transformation matrix of the moving frame with respect to the fixed base, and ~r \u00bc x y z\u00bd T \u00f03:2\u00de is the variable column matrix of the relative position vector of P, evaluated in the space of the moving frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure2-1.png", + "caption": "Fig. 2 (a) Illustration of the deformation of an annular plate; (b) illustration of an infinitesimal element", + "texts": [ + " Concentrated load Fr is applied to rigid shaft center O, and the displacement of the inner surface is equal to the displacement of the rigid shaft. Assuming that the center O of the annular plate is O# after deformation, the displacement of the center to a previous shape in the radial direction is \u03b4 \u00bc OO#. Under force Fr, the annular plate produces tensile and shear deformation. According to the energy conservation law, the work W done by external force Fr is transferred to elastic deformation energy U. Elastic energy U has two components, namely, tensile deformation energy U1 and shear deformation energy U2, where U=U1+U2. As shown in Fig. 2, we consider an infinitesimal element with angular displacement dq at angle q. The side lengths of the trapezium shape are a\u2012b, adq, and bdq. The tensile deformation is t \u00bc \u03b4cos , which can be derived as t \u00bc ! a b dFdr Ehrd \u00bc dF Ehd ! a b dr r \u00bc \u00f0lna \u2013 lnb\u00dedF Ehd : (10) The expression indicates that tensile deformation t is proportional to tension dF. Hence, dU1 \u00bc ! t 0 dFdt \u00bc 1 2 dFt \u00bc Ehd 2\u00f0lna \u2013 lnb\u00det 2 \u00bc Eh\u03b42cos2 d 2\u00f0lna \u2013 lnb\u00de : (11) By integrating Eq. (11) from 0 to 2\u03c0, we obtain tensile deformation energy U1 as follows: U 1 \u00bc " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002762_icelmach.2018.8506797-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002762_icelmach.2018.8506797-Figure1-1.png", + "caption": "Fig. 1. Cross-section of the prototype HEPMM.", + "texts": [ + " The operating speed range can be further extended by Method-II and Method-III. In addition, Method-III can provide a higher efficiency in flux-weakening region than Method-II since the copper loss of field winding can be decreased in proportion to the reduction of field excitation current. Those flux-weakening control methods are verified by experimental results. Index Terms\u2014Efficiency improvement, flux-weakening, hybrid excitation, optimal. I. INTRODUCTION ybrid-excited permanent magnet machines (HEPMMs), as shown in Fig. 1, consist of two excitation sources, i.e. permanent magnets (PMs) and field windings. HEPMMs provide an extra flexibility to adjust the flux-linkage by the field excitation current. Consequently, higher torque at low speed and wider operating speed range, as well as higher efficiency over wider operation region can be provided [1]- [5]. Moreover, due to windings and PMs are located on the stator and a simple robust salient pole rotor, high-speed operation and cooling capability are relatively easier for such machines [1]-[7]", + " In Method-III, the maximum efficiency condition is analyzed based on the differentiation method in order to determine the optimal field excitation current. All flux-weakening control methods are validated by experimental results. In this section, the novel fault-tolerant prototype hybridexcited machine and three flux-weakening control methods, which are classified based on the utilization of weakening currents, i.e. utilizing field excitation current only, utilizing armature current only, and optimal method, will be explained as follows: Fig. 1 shows the simplified structure of prototype HEPMM which consists of 12-stator/10-rotor poles. In this topology, four armature coils in each phase (A, B, and C) and twelve field coils are separately connected in series, and those coils are wound on each of the stator poles. The slot openings between two stator poles are replaced by the PMs. Since there has no magnet or excitation coils on the rotor, the rotor structure is simple. The phase flux-linkage for a specific field excitation current which is obtained by 2D finite element (FE) analysis is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003931_b978-0-444-64156-4.00001-5-Figure1.2-1.png", + "caption": "Figure 1.2 A line contact formed by a pair of cylinders. The pressure distribution in dry contact on the left and the film shape for the piezoviscous-elastic lubrication regime on the right.", + "texts": [ + " This is the important full elastohydrodynamic regime that is commonly just known as elastohydrodynamics or elastohydrodynamic lubrication (EHL) or elastohydrodynamics (EHD) and is operating in rolling element bearings and gears lubricated by viscous organic liquids. Johnson [3] identified the parameters that should characterize the particular regime of lubrication, the dimensionless viscosity and elasticity parameters, gV and gE, respectively. For line contact, these may be written as g2V 5 8\u03c03\u03b1 2p 6 H R E03u\u03bc0 (1.1) g2E 5 4\u03c02p 4 H R E03u\u03bc0 (1.2) Here, \u03bc0 is the viscosity of the liquid at atmospheric pressure, \u03bc\u00f0p5 0\u00de, and \u03b1 is a pressure viscosity coefficient defined by 1 \u03b1 5 pai 5 \u00f0N 0 \u03bc p5 0\u00f0 \u00de \u03bc p\u00f0 \u00de dp (1.3) Fig. 1.2, showing two cylinders in contact, will be helpful in defining the other parameters. The composite radius is R5 1 \u00f01=R1\u00de1 \u00f01=R2\u00de (1.4) and the reduced modulus of the solid bodies is E0 5 2 12 \u03bd12\u00f0 \u00de=E1 1 12 \u03bd22\u00f0 \u00de=E2 (1.5) The maximum contact pressure for a dry contact of width 2b, as shown in left portion of Fig. 1.2, is pH and this pressure will not necessarily be the maximum pressure in the liquid film shown in the right-hand portion of Fig. 1.2. The shape of the film will differ for the various regimes; the film shown in the right of Fig. 1.2 is the case of piezoviscous-elastic lubrication. Johnson constructed a chart in gV and gE coordinates to delineate the four regimes of line contact lubrication. Hooke [4] adjusted the boundaries of Johnson\u2019s regime chart to fill in some missing portions. The boundaries in such charts are generally constructed by finding the relationships between gV and gE that make the theoretical solutions for film thickness equal for the two regimes at the boundary. Winer and Cheng [5] have shown the boundaries as \u201cfuzzy\u201d in Fig", + "8) The first term on the right, hs, is a constant to be determined, the second is an approximation to the shape of the gap between the cylinders near the conjunction, and the last term is the elastic deflection resulting from the pressure distribution, p(x). The deformation at any position in the contact depends upon the pressure everywhere. For dry contact, the pressure distribution is the theoretical Hertz pressure distribution, p\u00f0x\u00de5 pH 12 x b 2 \u00f01=2\u00de (1.9) The Hertz contact half-width, b, shown in Fig. 1.2, is for dry contact b5 4pHR E0 (1.10) and substitution of the Hertz pressure distribution (1.9) into the film shape Eq. (1.8) and integrating yields a uniform thickness that can be made uniformly zero by the adjustment of hs as expected. The Newtonian assumption has been the basis for engineering film thickness formulas. The Reynolds equation that relates local pressure to the film shape, the Newtonian viscosity, \u03bc, velocity, u, and density, \u03c1, is for steady conditions, d dx \u03c1h3 \u03bc dp dx 5 12u d\u00f0\u03c1h\u00de dx (1", + " hm5 1:26\u03b1 0:6\u00f0\u03bc0u\u00de0:7E00:16R0:3p 20:26 H (1.14) Dowson and Toyoda [9] found an expression for the incompressible central film thickness. hc 5 2:55\u03b1 0:56\u00f0\u03bc0u\u00de0:69E00:07R0:31p 20:20 H (1.15) They found that given the same conditions, the minimum film thickness is about 80% of the central value. Much of the contact area can be characterized by a nearly uniform film thickness with the small variation resulting from compressibility. The minimum film thickness occurs at a restriction at the liquid exit to the contact as in Fig. 1.2. The very large negative pressure gradient at the exit, dp=dx52 2pH=b, would generate a substantial flow out of the contact at this location. The appearance of the exit restriction avoids continuity of flow problems. A local spike in the pressure occurs just upstream of this restriction as a consequence of the pressure dependence of viscosity. Hamrock [6] found for the compressible central film thickness, hc 5 2:154\u03b1 0:47\u00f0\u03bc0u\u00de0:692E00:110R0:308p 20:332 H (1.16) The two-dimensional form of the steady Reynolds equation required for a point contact solution is @ @x \u03c1h3 \u03bc @p @x 1 @ @y \u03c1h3 \u03bc @p @y 5 12u @\u00f0\u03c1h\u00de @x (1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001642_j.ejc.2018.02.023-Figure14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001642_j.ejc.2018.02.023-Figure14-1.png", + "caption": "Fig. 14. S is foldable into E.", + "texts": [ + " Then it is easily verified, by using an equi-rotational transformation for Conway tile N, that N is foldable into an envelope if there exists P such that P is divided into two identical right triangles by its diagonal (Fig. 10(a)\u2013(c)). We call such a Conway tile N bi-right. Consider a Conway tile N1 in Fig. 11. Although N1 is not bi-right, but it can be folded into an envelope (Fig. 12). This fact can be explained as follows: 1. N1 is equi-rotational into a parallelogram stripS by hinging pieces at three points of its 4-base {v1, v2, v3, v4} (Fig. 13). 2. S can be folded into an envelope E whose corners are v1, v2, v3 and v4 (Fig. 14). 3. By Corollary 1, N1 is also foldable into the same envelope E. The paper [1] studies how many ways there are for a given strip to be folded into different rectangle dihedra. On the other hand, in this paper we focus on positioning 4-bases (4 corners of dihedra) of foldable strips into rectangle dihedra (Lemmas 1 and 2) and finding more kinds of figures which are foldable into rectangle dihedra (Theorems 3 and 4). Lemma 1. A parallel strip S1 can be folded into an envelope E if S1 can be decomposed into 4n(n + 1) identical right triangle Ts, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002924_978-3-030-03451-1_23-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002924_978-3-030-03451-1_23-Figure2-1.png", + "caption": "Fig. 2. Sampling-based grasp planning", + "texts": [ + " The three degrees of freedom in the order symmetry, translation and tolerance are used and discretized using a Van der Corput sequence [20]. The number of samples is iteratively increased until a final gripping point (sample) has been determined. Figure 1 shows different approaches or gripping positions for a symmetrical object. The different stages in the figure represent the usage of more degrees of freedom during the sampling developed in this concept. The first stage uses rotatory symmetry, the second uses additionally vertical translation while the last stage changes the approach angle in a defined tolerance range (Fig. 2). In the next step, the individual samples of the various combinations are checked in a 3D simulation environment for collision avoidance and accessibility by the available robot. In doing so, defined approach vectors, collisions between gripper, component, robot and environment are taken into account in the description. In the case of mechanical grippers, it is also checked whether a grasp could be achieved when reaching the end position. The results of the gripper planning are individual points in space such as the gripper\u2019s or workpiece\u2019s pick-up position as well as approach and drop-off positions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001986_s0025654418010065-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001986_s0025654418010065-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " Object of Study Constructively, a rotary parachute can be realized on the basis of a conventional parachute of the \u201cflat circle in cutting\u201d type, whose sectoral cloths are deployed around their axes of symmetry by shortening a part of parachute lines. We will consider as a model a solid body of the following construction. Some number n (n > 2) of identical flat blades with area S/n are installed on consoles orthogonal to the vertical axis O1O and lying in one plane. Each blade is a thin symmetrical plate, stretched along the console. All blades are rotated around their consoles to the same setting angle \u03d5 (Fig. 1a), counted from a vertical. All this rigid structure can rotate as one solid body around the axis O1O with angular velocity \u03a9. We denote by V the vertical velocity of the rotational body (point O) (V > 0, if rotational body moves upward). The working mode of a rotochute is its descent along a vertically oriented axis O1O and rotation around this axis. *e-mail: samson@imec.msu.ru 51 1.2. Model of Aerodynamic Impact It is assumed that the air flow acts only on the blades, the resulting aerodynamic forces are applied at the points Ci of the blades at a distance r from the axis of rotation (Fig. 1 a). Fig. 1 b presents the calculation scheme for rotational body descent mode and the following designations are introduced: vector U (U = r\u03a9) is the horizontal component of velocity of point C, vector W is the speed of point C relative to air (airspeed), \u03b1 is the angle of attack, vectors L and D are lifting aerodynamic force and drag force, respectively. An auxiliary angle \u03b8 is introduced to simplify further presentation; obviously, \u03d5= \u03b1 + \u03b8. In accordance with the quasistatic hypothesis, values L and D of aerodynamic forces have the forms L = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002820_detc2018-85656-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002820_detc2018-85656-Figure1-1.png", + "caption": "Fig. 1 Mating process of a spur gear pair with friction forces", + "texts": [ + " The consideration of the gearbox casing in the dynamic model not only enables the study of the vibration features of the whole transmission comprehensively but also provides a measurable signal for direct comparison with vibration signals obtained from the outside of the gearbox casing. In light of the above, the goal of this paper is to propose a new dynamic model with consideration of the gearbox casing in order to simulate the vibration behavior of a gear transmission. The proposed model incorporates TVMS, time-varying load sharing ratio, dynamical tooth contact damping ratios, friction forces and friction moments. Experimental tests are performed to validate the proposed dynamic model under different speed conditions. Fig. 1 shows the mating process of a spur gear pair. where 0 is the pressure angle, 1 and 2 are rotational speeds. 1 2B B is the line of action, 1 2PP is the actual gear tooth contact line. pgF is the total dynamic mesh force. 1gF and 1pF are friction forces of the first gear pair, 2gF and 2pF represent the friction forces of the second gear pair. 2b is the width of the lubrication oil under Elastohydrodynamic Lubrication (EHL) conditions. 1pu , 1gu are surface moving speeds a contact point along the off-line of action [30]. 1pr , 2pr , 1gr , 2gr are equivalent radii of the mating teeth pair given by ( )1 1 0 tan( ) , 1,2 ( )sin pi bp p P gi pp pg pi r r B O P i r r r r = + = = + \u2212 , (1) in which ( )1 1 1 arccos p bp p B O P R O P = with ( ) 2 2 11 2 cos p ag p g ag p g p g O P R O O R O O O O P= + \u2212 and ( ) 1 0 arccos p g bg ag O O P R R = \u2212 [31] (see Fig. 1), and p is the rotational angle of pinion. Based on [32], the dynamical central film thickness and minimum film thickness of the contact point can be expressed as ( ) ( ) 0.135 0.705 0.556 1.222 0.223 0.229 0.748 0.842 0.077 0.716 0.695 min min 1.120 0.185 0.312 0.809 0.977 2.691 1 0.2 1.652 1 0.026 c Hc H h K W U G R V W U G h K W U G R V W U G \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 = + = + , (2) where W is the dimensionless load, U is dimensionless speed, G is dimensionless material coefficient, V is dimensionless hardness, R = is dimensionless surface roughness[32]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002495_ijaec.2018100103-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002495_ijaec.2018100103-Figure1-1.png", + "caption": "Figure 1. View of SWIPR system", + "texts": [], + "surrounding_texts": [ + "A\ufeffview\ufeffof\ufeffSWIPR\ufeffis\ufeffshown\ufeffin\ufeffFigure\ufeff1.\ufeffThe\ufeffaim\ufeffis\ufeffto\ufeffmaintain\ufeffthe\ufeffupright\ufeffposition\ufeffof\ufeffrobot\ufeffchassis\ufeff while\ufeffthe\ufeffwheel\ufeffis\ufefffree\ufeffto\ufeffmove\ufeffin\ufeffhorizontal\ufeffdirection\ufeffalong\ufeffwith\ufeff360\u00b0\ufeffrotation. The\ufeffmathematical\ufeffequations\ufefffor\ufeffSWIPR\ufeffwere\ufeffderived\ufeffusing\ufeffNewton\u2019s\ufeffsecond\ufefflaw\ufeffof\ufeffmotion.\ufeffThe\ufeff governing\ufeffmathematical\ufeffequations\ufeffare\ufeffgiven\ufeffbelow. x M L g M ML J M L ML J T M ML J M L R e e = \u2212 +( )\u2212 + +( ) +( )\u2212 2 2 2 2 2 2 2 2 2 \u03b8 \ufeff (1) \u03b8 \u03b8 = +( )\u2212 \u2212 ( ) +( )\u2212 MgL M M ML J M L ML T M ML J M L R e e e 2 2 2 2 2 2 \ufeff (2) where,\ufeffM W M J Re = + + 2 In\ufeffthe\ufeffabove\ufeffequations\ufeff\u03b8\ufeffand\ufeffx,\ufeffrepresents\ufeffthe\ufeffangular\ufeffdisplacement\ufeffand\ufeffhorizontal\ufeffdisplacement\ufeff of\ufeffSWIPR.\ufeffThe\ufeffvalues\ufeffof\ufeffvarious\ufeffparameters\ufeffconsidered\ufeffin\ufeffthe\ufeffstudy\ufeffare\ufeffshown\ufeffin\ufeffTable\ufeff1." + ] + }, + { + "image_filename": "designv11_92_0002865_icnp.2018.00043-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002865_icnp.2018.00043-Figure5-1.png", + "caption": "Fig. 5. A sample trace of finding the positions of an UAV.", + "texts": [ + " This process is repeated until an equilibrium state is reached where a particle has zero velocity. In this paper, we assume that when the movement of an UAV is less than a certain predefined threshold, the static equilibrium for that UAV is achieved. In the final step of the algorithm, we apply the outcome of the Utility Function to judge whether we should put an UAV in this area. If the value of the Utility Function is larger, we add the corresponding UAV to the area, otherwise not. In order to illustrate the equilibrium point finding mechanism (i.e., EMech) of an UAV, we plot Figure 5 that shows a sample trajectory of an UAV in the process of finding its optimal position from its initial point. In this figure, 10 users are non-uniformly distributed and they are marked by black dots. Each UAV chooses its initial position randomly that is denoted by a blue dot. The movement of the UAV is decided by the net force from the users and the temporary stops are indicated by the stars. The process is repeated until it reaches the equilibrium state and the final position is specified by a red dot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002004_jifs-169540-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002004_jifs-169540-Figure8-1.png", + "caption": "Fig. 8. Experimental setup of a DDS system.", + "texts": [ + " The mathmatical expressions of these parameters are listed in Table 5, where x(t) represents the original gear vibration signal, \u03bc is the average value of the signal x(t) and xp is the peak value of the signal. The diagnostic results are shown in Table 6, where 7 data samples are misclassified, leading to 91.25% classification accuracy, which is lower than that of reccurrence network-based approach. The second experimental study was conducted on a Drivetrain Dynamics Simulator (DDS) platform for characterizing different types of gear faults, as shown in Fig. 8. Table 7 lists different gear faults tested in this study. Vibration signals are acquired with 1024 Hz sampling rate and 512 s sampling window when the simulator is running at 30 Hz rotating speed. Figure 9 illustrates the waveforms of the gearbox vibration signals under four different working conditions. Using the FNN algorithm and mutual information, the embedding dimension and time delay are selected as 6 and 2, respectively. Threshold for constructing recurrence matrix is set as 0.4. Through recurrence network theory, the features , \u03c1, C, As , T, HSE and IR are also extracted from the vibration signals, and Table 8 lists sample features from each condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002082_1.4040813-Figure15-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002082_1.4040813-Figure15-1.png", + "caption": "Fig. 15 Pressure profiles versus bearing load, 12,000 rpm, 100% flow rate", + "texts": [ + " In both the flooded and starved models, increased bearing load results in an increase in both stiffness and damping over all speeds. Increases in damping ratio with load in both models suggest that the increase in stiffness with load is driving these trends. For each load, the starved model predicts lower bearing stiffness and damping than the flooded model across all speeds. Higher predicted damping ratios and lower damped natural frequencies in the starved model again suggest that the differences in stiffness values are dominating the trends seen in the resulting modal properties. Figure 15 shows the predicted pressure profiles of the starved bearing model under both load conditions at the nominal flow rate and 12,000 rpm. Under the increased bearing load, the predicted pressure on the bottom two pads increases significantly, as expected. Also, upper pads that are partially or lightly loaded under the nominal load condition are completely unloaded under the increased load. The net result is the increase in bearing stiffness and damping with load seen in Figs. 13 and 14. For both load conditions, the experimentally identified damped natural frequencies fall very close to the starved model predictions, while the identified damping ratios are consistently lower than starved model predictions and fall closer to the flooded model predictions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000992_ical.2008.4636350-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000992_ical.2008.4636350-Figure3-1.png", + "caption": "Fig. 3 Sketch diagram of exciter and sensors position", + "texts": [ + " The two exciters are positioned perpendicular to each other at 45\u00b0 to the horizontal, pointing to the geometric center of the test bearing. Each exciter is manipulated by an individual controller and driver and can generate sinusoidal force which frequency and amplitude are adjustable. C. Data Acquisition The tested data in the test consist of the bearing radial displacements, journal rotating speed, frictional force between the journal and the test bearing, and the dynamic forces of the exciters. The positions of the testing sensors are indicated in Fig.3. Four electric eddy current sensors with a sensitivity of 8V/mm are used to measure the relative displacement between the bearing and the journal in the vertical and horizontal directions. They are installed on both ends of the test bearing sleeve. The rotating speed of the journal is measured by an OMRON E3S-CL photoelectric transducer which is mounted on the test platform near the output end of the motor and can produce a square wave signal in each revolution of the journal. A pressure sensor (Fig.3-7, CL-YB-61) connected to the test bearing sleeve by a torque rail is used to measure the frictional force between the surface of the journal and the test bearing. This frictional force can be used to identify the lift off speed of the test bearing [11]. The other two pressure sensors measure the dynamic forces of the exciters. They are mounted in the thin-walled tubular connecters which connect the exciters to the bearing sleeve. In experiments, each sensor is equipped with a separate amplifier to reach high metrical accuracy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001891_irec.2018.8362457-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001891_irec.2018.8362457-Figure3-1.png", + "caption": "Fig. 3. Points where the flux density vector and B-H relation have been computed.", + "texts": [ + " The obtained values are then used as input of the applied magneto-dynamic vector hysteresis model. Therefore, a two dimensional finite element analyses is performed iteratively until the attainment of the convergence. IV. Simulation results In this study, we focus on the investigation of magnetic field and the modeling of the dynamic hysteresis behavior in the studied machine. Because of the stator core is of main interest, local computation has been performed at particular elements belonging to the stator core. The chosen elements are marked by points illustrated in Fig. 3. The variations of the magnetic flux density in those particular elements are illustrated in Figs. 4, 5, 6, 7, 8. From Figs. 4 and 5, we notice that the flux density variation in the point a which is near to the outer of the stator yoke is pulsating along the tangential components. On the contrary, the flux density variation in point b belonging to the stator yoke is approximately rotating. Referring to Figs. 6, 7, 8, we can easily remark that the flux density variation, in the stator tooth tips and the tooth end is not entirely alternating" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002432_978-3-319-99522-9_8-Figure8.1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002432_978-3-319-99522-9_8-Figure8.1-1.png", + "caption": "Fig. 8.1 The 3-PRR planar parallel manipulator", + "texts": [ + " Provided with a closed-loop structure, the planar parallel 3-PRR manipulator is a special symmetrical mechanism composed of three planar kinematical chains with identical topology, all connecting the fixed base to the moving platform. The points A0;B0;C0 represent the summits of a fixed triangular base and three moving revolute joints A3;B3;C3 define the geometry of the moving platform. Every leg consists of two links, with one prismatic joint and two revolute joints. The parallel mechanism with seven links \u00f0Tk; k \u00bc 1; 2; . . .; 7\u00de consists of three prismatic and six revolute joints (Fig. 8.1). Whereas the kinematics degrees of freedom, or mobility, of a serial chain robot can be obtained as the sum of the degrees of freedom of each of the joints, the situation for parallel robots or closed chains in general is somewhat more complex, since only a subset of joints can be independently actuated. The Gr\u00fcbler-Kutzbach equation predicts that the device has certainly three degrees of freedom [1]. In the actuation scheme PRR each prismatic joint is an actively controlled prismatic cylinder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000814_s11029-009-9095-4-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000814_s11029-009-9095-4-Figure2-1.png", + "caption": "Fig. 2. Trans for ma tion of the bil let into the sec tion of a cir cu lar to rus.", + "texts": [ + " 3370191-5665/09/4504-0337 \u00a9 2009 Springer Sci ence+Busi ness Me dia, Inc. Cen tral Re search In sti tute of Spe cial Me chan i cal En gi neer ing, Khotkovo, Mos cow Re gion, Rus sia. Trans lated from Mekhanika Kompozitnykh Materialov, Vol. 45, No. 4, pp. 497-506, July-Au gust, 2009. Orig i nal ar ti cle sub mit ted Sep tem ber 1, 2008; re vi sion sub mit ted April 3, 2009. Me chanics of Com pos ite Ma te rials, Vol. 45, No. 4, 2009 For a cir cu lar to rus, as a shell of rev o lu tion, the geo met ric re la tions (Fig. 2) ds Rd= J tan b, dl Rd = J cos b , d ds r a = (2) are valid, where r R t= +( cos )J and t r R= 0 . The last equa tion in (2) can be pre sented as d d t a b= + tan J Jcos . (3) By def i ni tion, the wind ing an gle b is the an gle be tween the me rid ian (a cir cum fer ence of ra dius R) and the filament. Let us es tab lish the con di tions and re la tions for trans for ma tion of the cy lin dri cal shell-bil let into the sec tion of a cir cu - lar to rus. It is ob vi ous that they are de ter mined by char ac ter is tics of the man drel on which the bil let is wound. In this study, we will con sider man drels in the form of a cor ru gated pipe or a spring able to de form in ten sion-com pres sion along the generatrix, but which is rigid in the circumferential direction. Let there be a generatrix on the man drel whose length re mains un changed dur ing the de for ma tion pro cess. Then, upon bend ing the bil let, to this generatrix there cor re sponds the radius r R tH H= +( cos )J , (4) on the branch (see Fig. 2), de ter min ing the neu tral line, along which the de for ma tion of branch generatrix is e = 0. Ow ing to sym me try, in the gen eral case, there are two such generatrices, cor re spond ing to \u00b1J H . To the neu tral line there cor re sponds the ec cen tric ity e R H= -cos J rel a tive to the midline r r= 0 of the branch. Since the length of the el e ment of the man drel generatrix (and hence of the bil let generatrix) along the neu tral line re - mains the same dur ing the trans for ma tion pro cess, the relation ds r dH3 = a, (5) is valid and, ac cord ing to Eqs", + " The sec ond re stric tion is de ter mined by the fact that the ini tial geo detic tra jec tory of wind ing of the shell bil let be comes nongeodetic on the to rus, and for re al iza tion of the trans for ma tion in the sec ond vari ant, the fil a ment must be in equi - lib rium owing to friction forces. Ac cord ing to the Meusnier the o rem [5], the cur va ture of the fil a ment is k kn= cos q , where q is the an gle of geo detic de - vi a tion. The nor mal cur va ture of the filament is k k kn = +1 2 2 2cos sinb b. (16) Here, k1 and k2 are the main cur va tures of the sur face of a cir cu lar to rus. For fur ther cal cu la tions, we will in tro duce the an gle J J1 2= +p as a vari able (see Fig. 2). Then, the equa tion of the cir cu lar to rus takes the form r R t= +( sin )J 1 . Ac cord ingly, 30\u00b0 and t = 5 ( ); 10 (- - -). e R = 0 (1, 4); 0.5 (2, 5); 1 (3, 6). k R 1 1 = , k r R t 2 1 1 1 = = + sin sin ( sin ) J J J . (17) For the fil a ment to be long to the sur face, the con di tion k \u00b3 0 must be ful filled, which is equiv a lent to the re stric tion kn \u00b3 0. Then, us ing Eqs. (16) and (17), we obtain 1 01 1 2+ + \u00b3 sin sin J Jt tan b . (18) Re stric tion (18), with ac count of Eq. (12), can be pre sented as ( cos ) sin ( sin )t tH+ + + \u00b3J J J2 1 2 3 1 0tan b " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002355_978-981-13-0629-7_5-Figure5.10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002355_978-981-13-0629-7_5-Figure5.10-1.png", + "caption": "Fig. 5.10 Prototypes and test platform of 12-slot/10-pole Sine+3rd-shaped PM machine", + "texts": [ + "9 shows flux density distributions for three machines with optimal k under c = 0.5 and c = 0.8 when flux linkage of phase A is maximum. It can be seen that the maximum flux density in stator lamination is around 1.5 T, which is below the knee point of lamination steel B-H curve, 1.6 T. The saturation level of the machines is low. 5.5 Experimental Verification The 12-slot/10-pole SPM machine with sine+3rd-shaped rotor has been prototyped and tested for validation. The stator and the sine+3rd-shaped rotor are pictured in Fig. 5.10. Its stator outer diameter (Dso) and active axial length (Lef) are 90 and 50 mm respectively. It is designed with the optimal split ratio (k = 0.6, Dsi = 27 mm) with flux density ratio is 0.5. The test platform is shown in Fig. 5.11. Figure 5.11a shows variation of the FE predicted and measured cogging torques with rotor position. Although the peak value of cogging torque is slightly higher than the FE predicted one probably because of the manufacturing tolerances and assembly deficiencies. The agreements between the FE predicted and measured of cogging torque are good" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002046_978-3-319-79005-3_17-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002046_978-3-319-79005-3_17-Figure1-1.png", + "caption": "Fig. 1 Shell181, geometry, nodes and coordinates", + "texts": [ + " There are many elements in the libraries of each program based on the FEM that adapt to the type of analysis to study, which in the case of ANSYS are the element types beam, solid, shell, among others. SHELL181 is suitable for analyzing thin or moderately thick structures. It is an element of four nodes with six degrees of freedom in each node: translations in the X, Y, Z, and rotations on the X, Y, and Z axes. The option of triangular elements should only be used as filler elements in the mesh generation. Figure 1 shows the geometry, location of the nodes, and the coordinate system for this. The element is defined by four nodes: I, J, K, and L. Temperatures can be introduced as loads to the element at the 1\u20138 corners. The first corner temperature T1 is set by default to the uniform temperature (TUNIF). The other temperatures, if not specified acquire the temperature T1. This element has the characteristic of generating layers with orientations that allows simulating the printed specimens in ABS. 3D printing has gained a lot of popularity in recent years, owing to the interest consumers have had for low-cost 3D printers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001483_978-3-319-72730-1_17-Figure17.2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001483_978-3-319-72730-1_17-Figure17.2-1.png", + "caption": "Fig. 17.2 Field and current layer orientation in a DC machine", + "texts": [ + " Recent scientific papers have shown, however, that appropriate control methods may make V/f control of synchronous motors quite feasible and practical. Still, vector control or similar methods are much more common for synchronous machine drives. The torque of an idealDC commutatormachine (with a brush axis electrically orthogonal to the flux axis and no armature reaction) is proportional to the product of flux and armature current (seeEq.14.1). This is because the symmetry axis of the armature current layer is co-incident with the field flux axis, and the torque is at a maximum for a given armature current and flux (Fig. 17.2). At the same time, if there is no armature reaction,1 a variation of the armature current will not affect the flux and transients in the armature current will not influence the torque-producing flux. As a consequence, there will be no additional transients. In rotating field machines, the torque results from the interaction of a rotating sinusoidal field distribution b(x, t) and a rotating sinusoidal current layer a(x, t), as was explained in Chap. 3 of Part 1. As the speeds of both field distribution and 1For a brush axis in the neutral position, this requires a compensation winding or the absence of saturation-induced armature reaction; however, a brush axis which is not in the neutral position may result in armature reaction for non-rated operating conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001718_icit.2018.8352487-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001718_icit.2018.8352487-Figure5-1.png", + "caption": "Fig. 5. Prototype for experiment", + "texts": [ + " The biggest obstacle in indoor environment is assumed the stair, the position damping is investigated while the robot flies over the same distance as the horizontal length or the stair. The stair in the experiment is the minimum stair having 13 steps with 0.15 tread being regulated by Japanese building law. The horizontal length of this stair is 1.95 [m]. The initial velocity is set to 0.5[m/s] and flying horizontal distance is set 2.0 [m], and then the flying time is calculated as follows by using uniform motion model: t = x v0 = 2.0 0.5 = 4.0[s] (15) The robot is used in this experiment is shown in Fig. 5. The robot is combined of Parrot AR.Drone 2.0 and 3 wheels robot. Each mobile mechanism take charge of spatial moving and moving on ground. This robot is bigger than the proposed prototype shown in Fig. 1. Therefore the coefficient of air friction and projected area of Fig. 5 is bigger than the prototype (Fig. 1). Thus if the position damping can be neglected using AR Drone, the position damping can also be neglected in various size of mobile robot. Therefore the robot shown in Fig. 5 is adopted as benchmark of worst case of air friction. The distance of the robot is measured by kinect for windows v2. The runway(0.5[m]) is prepared for maintaining initial velocity 0.5[m/s] at take off. When the robot flies, the rotors never generate thrust force for moving forward, the robot moves by using only inertial force made by wheel motor on the runway. The motion and result are shown in Fig. 6, Fig. 7. In Fig. 7, the robot moves on runway to reach the initial velocity by 1[s]. It flies by 1[s] to 5[s]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000824_2009-01-0610-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000824_2009-01-0610-Figure4-1.png", + "caption": "Figure 4- F s acting on the $%& rolling elements orce", + "texts": [ + "/ 012 3# 45!673892/ f ure 2, the contact angles can be expressed as: (4) :;'*+, -./4-./ 01? 3# 45!67389?/ (6) :;'*=, > ,,: ',. :: :. ~:=::~:~ ~ '~~;~\" L: : : ]L : ::Il, :: : iii i i i~ i -20 -- --- - -~ --- ----:--- ---- -:--- ----~- -- ----~- -- ----~ - 1-- ---: -:', -- --~- --- -- : : : : : : = : : I I I I I \" , I , \" \" I I I I I \" -300:--1J.:0-~20,------::3~0--4-,1:0,----::50':------:6J.:0---:l70,------::8~0 --:190 Time t(s) Fig.1 3 Thrust and velocity calculation curves -300!-- 1..':0- --:-'20,------:3:'o0- ---'40,------::5LO--6::':0'--70L--S.L0---:-'gO Time t( s) Fig.14 Thrustand velocity experiment curves , , ,, , . , 10 ------- ~ -----f-------f------- f-------t------- -------i-- ----i-----, I \" \" , I \" \", , , , , , . I , I I , I , I o - - - - - - - ~ - - (l )Velocily v2(m/s)(6:1) ------ -i ------:--- ---- - ---- ! _ (2)Thrust Fe(N) (1:5) ! ' ' , , , \u2022 , \u2022 \u2022 , I , -10 ----- -- ~ - -- - - -- ~ ------ - ~ --- ---- ~ - ------j---- ---j- --- --i---, , I , , I , \u2022 , I , I I , , , , , , I , -20 -------~ ------- ~ -------~ -------~ -------j-------!-------:-------j-----, , I , , I , , \u2022 , I , , I , , , , I , , I , , , I I , I I , I \u2022 , , , , I , , 40 r--..,-----,-----,-----,--,--,----,-----,-----, 40 r--,--,---,---,----.-----,----,-----,-- and q axis current, which are modified by three PI regulators'\". Fig.13 is thrust and velocity simulation curves, Fig.14 is thrust and velocity experiment curves, which effectively validate each other. 12 SLIM 10 50Hz \\ ..~ . ';,\\ v \\. \\ \\ \\ 40Hz30Hz 4 6 S Vel oc i t y V2(m!s) 20Hz 10 100 r-----.-------,----,-----.-------,------, 90 10Hz ~ v~ u, I \", INV V, m; + V m, I \", I \" m, B, I , m, + I \", 2 / 3 I , + io, 30 20 Fig.11 Calculation and measure curves of thrust force under constantcurrentconstantfrequency In order to obtain high performance of linear motor vehicle, it is important to develop vector control for LIM . The basic idea behind the field oriented control is uncoupling the flux and the torque of an induct ion motor in order to achieve the torque respon se similar to that of a separately excited direct current machine. Similar is in case of linear induction motor except the torque has been replaced by thrust force and the rotational speed by linear speed . Thus, the field oriented control, can be adopted to decouple the dynamics of the thrust force and the secondary flux amplitude of the LIM [8 1 \u2022 VI. CO NCLUSION Fig.12Indirect orientation controlanalysisdiagram The aim of the field oriented control is to maintain constant the d-axis secondary flux and making null the q-axis secondary flux . The field oriented control utilizes rotor flux and torque feedback, obtained through direct sensing or an estimator. The rotor flux angle permits the decoupl ing of primary current into its force-producing component and its rotor-flux component. The dynamic model of the LIM is analyzed by using the d-q model of the equivalent electrical circuit with end effects included. Four coefficients are determined to express the effects that the LIM speed causes in the magnetization branch of the equivalent electrical circuit. The field oriented control for LIM can be analyzed in the same way that the traditional induction motor . Fig.12 is the indirect rotor field control scheme, which likes that of RIM control. There are three close loops, such as speed loop, d axis current This paper achieves linear induction motor new T model equivalent circuit considering end effects by using four coefficients, such as Kx(s) , Cx(s), Kr(s) , Cr(s) , to describe influence on mutual inductance and secondary resistance. Dynamic model and control strategies of secondary flux orientation for linear induction motor for high performance control have been presented. Simulation and experiment results are given to show that the control system has fast speed response and precise thrust characteristic." + ] + }, + { + "image_filename": "designv11_92_0000908_robio.2009.5420526-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000908_robio.2009.5420526-Figure3-1.png", + "caption": "Fig. 3. The motion freedom of the remote robot(left) and the try-out processes (right).", + "texts": [ + " Constraint set, which specifies the capability of the remote robot. To find: Correct and optimized parameters for the remote robot to implement the expected action. The solving of this problem consists of two steps, namely, the virtual action planning process, and the training of the remote robot. The former results in an action instance working with the virtual target in the enhanced reality environment, the latter will result in a list of actions to be conducted by the remote robot in the real world. 1) Virtual Action Planning: Fig.3 gives a description of the structure of the robot, consisting of 6 steering gears, which forms a linkage with 6 freedoms. The coordinate frame of the robot was set as x-axis pointing towards outside the screen, y-axis towards the earth, and z-axis the direction defined with the right-hand rule. And the initial state of the robot was at vertical position with that all the rotation angles of the steering gears equal to zero. The virtual action planning could be formulated as to find the optimized parameter set for the virtual robot with motion constraints, which satisfies the following equation: min diff(Po, Tg) (1) where Po denotes the position of the target object; and Tg the position of the gripper, which specified by Tg = 5\u220f i=0 (Ti) (2) where Ti denotes the ith transformation matrix relative to joint i \u2212 1 in the linkages, and Ti = Trans(0, 0,\u2212Yi) \u00b7 Rot(X, \u03b8i) \u00b7 Trans(0, 0, Yi) (3) where Yi denotes the Y -coordinate of the ith joint at its initial position; and Trans() the translation transformation; Rot the rotation transformation to a designated axis; And T0 = Rot(X,\u2212\u03b80)", + " 2) Implementing the Real Action: For the parameter set obtained by the virtual path planning process, the following steps will be conducted in order to implement the real action at the remote site, i.e., Visual check of the simulation of action process driven by the path planning parameters. This is expected to find unexpected interferences with surrounding objects during the action, or incorrect parameters caused by computation tolerances, with the help of vision system; Checking the validity of the actions, as show in the right side of Fig.3. By means of sending the optimized operation parameter set to the controller of the remote robot, a \u201dteaching\u201d process will be implemented. Since the action parameters were directly written to the steering gear in our implementation, therefore, a pre-scheduled actuation sequence is required. The learning process of the robot is implemented by a tryout procedure. And the verified parameter set guarantees the validity of the action performed by the remote robot. We have implemented a testing platform with the follow- ing configurations: A robot with binocular vision system, which serves as the remote robot; A control server equipped with our enhanced reality working environment for robot telecontrol", + "003m within a distance of 1m. Real-time control of remote robot relies heavily on the network speed. A direct connection to the local server via serial ports may help to promote the response speed. At current stage, to implement real-time control of remote robot via public Internet connection is still difficult, due to the un-controllable network delays. The planning, learning and then acting procedure are simulate by the virtual robot in the enhanced environment, as shown in the right side of Fig.3. In this paper, we have presented a learning-based strategy for action planning in robot telecontrol, in which the sophisticated actions of the remote robot equipped with a binocular vision system could be pre-scheduled with a virtual robot at the control terminal. The remote robot will then be \u2019taught\u2019 with the validated action plan with a series of parameter sets obtained from try-outs, thus accurately implementing the dedicated actions assigned . The action planning process is implemented within a enhanced reality environment, in which both the virtual and the real robot will be displayed simultaneously for the purpose of being deeply immersed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-20-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001513_b978-0-12-814641-5.50004-4-Figure19-20-1.png", + "caption": "Fig. 19-20: Two quadrant -\u201clifting\u201d operation of a DC-machine Fig. 19-21 Fig. 19-22", + "texts": [ + " This is called quadrant 1. We disconnect bridge A and control bridge B as an inverter ( \u03b1 2 > 90\u00b0 with V 2 negative). Due to the mechanical inertia the machine can now operate as a generator. With an unchanged direction SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR 19.21 A lifting motor can also operate in two quadrant mode but then in quadrant 1 and 2 (or in quadrant 3 and 4!). Consider a lifting motor that lifts a load with a positive torque and rotates clockwise (fig. 19-21). This is motor mode in quadrant 1 (fig. 19-20) together with the rectifier operation of bridge A in fig. 19-18 (applied voltage V 1 ; armature current I 1 ; counter emf E 1 ) . If we reduce the firing of rectifier A so that V 1 decreases, then at a certain instant the motor torque will be insufficient to raise the load and the load will begin to drop. (fig. 19-22). This is known as load pulling. The direction of rotation of the motor reverses and since the flux has not changed, the polarity of E = k 1 \u0387 n \u0387 \u03a6 reverses from what it previously was. The emf is now E 2 (fig. 19-18) and if bridge A is controlled as an inverter the polarity of V 1 can reverse. With | V 1 |< E 2 the machine will deliver energy to the power grid. This is operation in the second quadrant as represented in fig. 19-20. The load L has to be sufficiently large for load pulling to occur otherwise operation in quadrant 2 is not possible. Note that one bridge (A or B) is sufficient for this type of two quadrant operation. 19.22 SPEED- AND (OR) TORQUE-CONTROL OF A DC-MOTOR We reconsider the rolling mill as presented in fig. 19-19 and to the right of fig. 19.23. Controlling bridge A as rectifier allows the motor to rotate clockwise. Bridge B as inverter allows the machine to brake (still in a clockwise direction!) until it is at stand still" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000036_detc2009-87370-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000036_detc2009-87370-Figure3-1.png", + "caption": "Figure 3. The collision area of a layer for a branching part", + "texts": [ + " It is given by: cos)( maxmax C qN C h z (9) where is the angle between the surface normal and the build direction, )(qN z is the z component of the unit normal ))(),(),(()( qNqNqNqN zyxS at a point q i bS S, S is the boundary surface of the part. If h calculated from Eq. (9) is smaller than mint , the layer is built with the thickness h. Then the extra material for this layer can be removed in machining process as shown in Fig. 2. Although the thin wall deposition techniques have these advantages, a potential problem exists in this technique. As shown in Fig. 3, the interference between the nozzle and the deposited layers may occur. A collision area in which the nozzle will collide with the deposited material, is illustrated in Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/idetc/cie2009/70870/ on 04/11/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 4 Copyright \u00a9 2009 by ASME Fig. 3. Collision checking is conducted using the ray tracing technique as follows. Let iL be the layer to be checked, the access directions of the nozzle to the overhang area for iL include all the surface normal of points on the side surface of iL , i lateralS . If a directed line, i pl , which is defined by (p, pn ), p i lateralS , intersects with any deposited layer Lj, ij , i.e., ijLl j i p (10) then collision between the nozzle and the deposited occurs. All the surface normals of the i lateralS need be tested" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002481_gt2018-76823-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002481_gt2018-76823-Figure6-1.png", + "caption": "FIGURE 6. OIL JET ANGULAR DIRECTION DEFINITION", + "texts": [ + " A multiple variables fitting of the collected data was performed in order to obtain an empirical formulation for friction torque in the form of the following equation: \u03c4corr = a \u00b7V b p \u00b7T c b \u00b7Pd (1) where \u03c4corr is the torque calculated following the correlative method (eq. 1). The coefficients a, b, c and d were calculated for both the shafts. Figure 5 shows the quality of the correlation predictability: predicted values are plotted against the experimental ones, al- most all the points are disposed in a straight line with unitary slope within an interval of \u00b1 5%. Oil Momentum Transfer Model Another source of losses to be taken into account is the momentum transfer between the gear and the oil impacting the teeth, this was calculated using the 0D model, see Figure 6, proposed and tested in a preceding work and described in [8]. The following simple equation 2 was hence exploited for this purpose: \u03c4 = \u03c0D2 j 4 \u03c1 jVjRp ( \u03c9Rp \u2212Vj sin(\u03b1) ) (2) where D j, \u03c1 j and Vj are the jet diameter, density and velocity respectively, \u03b1 is the injection angle expressed following the scheme in Figure 6. Test protocol Every test was performed by warming up the rig and reaching the maximum speed, when the torque measured was found to be steady the point was acquired and the velocity reduced. An example of the applied post-processing is reported in Figure 7, where the overall torque measured by the torque-meter is compared with the calculated bearings friction, the oil momentum transfer and the resulting windage torques. 5 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000649_13506501jet353-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000649_13506501jet353-Figure1-1.png", + "caption": "Fig. 1 Structure of a self-locking synchronizer system", + "texts": [ + " As it seems easier to coat plane surfaces with a friction \u2217Corresponding author: Department of Machine Elements, Engineering Design and Tribology, Leibniz University Hannover, Welfengarten 1A, Hannover 30167, Germany. email: poll@imkt.uni-hannover.de material, this paper investigates if the component tests can be substituted by simpler pin-on-disc type model tests. Therefore, the idea is to compare pin-on-disc type model tests and tests with real synchronizers regarding different aspects, i.e. friction characteristic and contact conditions. Figure 1 shows the components of a synchronising system in an exploded view. The cone is positively engaged with the gear. The synchro body, the synchro sleeve, and the synchro ring as well as the springs and the centering pieces of the pre-synchromesh unit are the components that rotate with the shaft. The synchro sleeve can be displaced axially during gear shift with the aid of the gearshift linkage in the direction of the gear. This presses the ring against the cone via the stop teeth. The difference in speed is reduced by friction between the conical friction surface of the ring and the cone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003691_robio.2018.8665156-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003691_robio.2018.8665156-Figure1-1.png", + "caption": "Fig. 1. Concept of this research.", + "texts": [], + "surrounding_texts": [ + "As shown in Fig. 2, our human\u2013robot collaborative system consists of \u2022 a high-speed robot hand (Section II-A), \u2022 an image-processing PC with a high-speed camera (Sec- tion II-B), \u2022 a real-time controller that receives the values of the state of the board (position and orientation) from the imageprocessing PC at 1 kHz and also controls the high-speed robot hand at 1 kHz, and \u2022 a board to be manipulated by the robot hand and a human subject (Section II-C)." + ] + }, + { + "image_filename": "designv11_92_0000213_s11665-008-9314-5-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000213_s11665-008-9314-5-Figure5-1.png", + "caption": "Fig. 5 Positions of hardness measurement. (a) On the steel bar and (b) microhardness profile, starting from the screw thread root in Alloy 1. It is also observed the deformation lines from the rolling process. Values are presented in Table 3", + "texts": [ + " From the microstructural evaluation, it was possible to verify the microstructure composition and the average GS following the ASTM E112-96 standard (Ref 6), as well as the decarburizing and discontinuities present in the steels. Measures of hardness were performed in samples removed from the bars after drawing and in the thread region in a 904\u2014Volume 18(7) October 2009 Journal of Materials Engineering and Performance Universal Hardness Equipment, Rockwell C scale, and using the load of 150 kgf. Four measurements were performed on the surface as shown in Fig. 5(a). For the evaluation of cold work effect on the thread root due to the rolling process, 10 Vickers hardness measurements were carried out starting just below the thread root until the center of the bar (Fig. 5b). To evaluate the general tensile strength of the bar after the drawing process, samples were prepared and carried out according to ASTM E8M-00 standard (Ref 7, 8), at speed testing of 1.0 mm/min in a EMIC DL10000 mechanical machine. The specimens geometry and dimensions are presented in Fig. 6. Three samples were tested for each studied steel. As the U-bolt is submitted to repeated load, fatigue failure is a major concern and a factor to be considered in the procedure for selection of its steel and fabrication process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000739_icept.2009.5270649-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000739_icept.2009.5270649-Figure2-1.png", + "caption": "Fig. 2 The mesh of thermo-sonic ball bonding", + "texts": [ + " The bond pad and bonder capillary are rigid materials, which are much harder than the FAB and assumed not to be deformed. The material parameters of FAB are listed in Table 1. Since the bonder capillary is considered as a rigid body, this leads to the rigid and elastic nonlinear contact pair under elastic condition and the rigid and elastoplastic nonlinear contact pair under elasto-plastic condition between FAB and bond pad. Mesh was created for the symmetric loading geometry using the FE code ANSYS, as shown in Fig.2. Supposed the friction coefficient of the contact interface is 0.4. The FEA model is composed of 7168 solid186 elements and 31265 nodes. This ensures sufficient resolution and accuracy of the results while maintaining a reasonable time needed for computation time. The elastic case and the elasto-plastic case have the same analysis condition except for the material of the FAB. 2.2. Analysis steps and solving In general, the real process of 1st ball solder point of thermo-sonic ball wire bonding can be refined as follows: Firstly, the FAB is formed by using high voltage discharge" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001073_epepemc.2008.4635363-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001073_epepemc.2008.4635363-Figure9-1.png", + "caption": "Fig. 9. Three phase SRM6-4", + "texts": [ + " 1, taken two times and with phase difference of superscript and the subscript ( 2 4 S RM ) the motor is a single phased SRM2-4 motor with 2 stator poles and 4 rotor poles. The simplified model can be seen in Fig. 8. The position feedback gives opportunity for 4 stable aligned 788 2008 13th International Power Electronics and Motion Control Conference (EPE-PEMC 2008) rotor locations. Evidently this description is very close to a single phased SRM6-4 motor type. What happens if the motor consists of three phases like the one shown in Fig. 9? Obviously the equation of the kind (2) will be tripled. difference, as in (3): ), 2 (),(),( 2 2 2 2 2 4 A S RA S RA S R UMUMUMPhA ), 32 (), 3 (), 3 ( 2 2 2 2 2 4 B S RB S RB S R UMUMUMPhB (3) ), 3 2 2 (), 3 2 (), 3 2 ( 2 2 2 2 2 4 C S RC S RC S R UMUMUMPhC After some transformations of (3) having in mind the periodical character of the neural network functions [0- , the following model of SRM6-4 can be obtained, as in (4): ), 6 3 (), 6 0 (),( 2 2 2 2 2 4 A S RA S RA S R UMUMUMPhA ), 6 5 (), 6 2 (), 3 ( 2 2 2 2 2 4 B S RB S RB S R UMUMUMPhB (4) ), 6 1 (), 6 4 (), 3 2 ( 2 2 2 2 2 4 C S RC S RC S R UMUMUMPhC The mathematical description (4) points out how to obtain the mathematical model of the SRM6-4 in Fig. 9, using only the description (1), i.e. 2 2 S RM . Superposing it 6 times with different initial rotor angles the system is ready to be used. There is also more compact way to describe the same system (5): (5) A graphical representation of (5) is pictured in Fig. 10. There are 12 stable aligned positions of the rotor in this model representation. In other words it is equal to 0.52[rad] step size. These steps can be seen by the stepper working mode of the motor, demonstrated in Fig.11. 5 )5()2( )4()1( )3()0( 0 2 2 ,, 2 4 ), 6 (),(,, BKii CKii AKii case i K S R CBAK K S R UiMUMCBPhA 2008 13th International Power Electronics and Motion Control Conference (EPE-PEMC 2008) 789 Phase currents and inductances are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002887_ihmsc.2018.00016-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002887_ihmsc.2018.00016-Figure1-1.png", + "caption": "Figure 1. The layout of the V/STOL.", + "texts": [ + " V/STOL PRINCIPLE ANALYSIS The V/STOL UAV discussed in this paper is based on the principle of propeller deflected slipstream for high lift, including a vertical propeller at the nose. The aircraft has a flying wing layout, with two control surfaces on one side, which are slipstream deflection control surfaces and attitude control surfaces . Four propellers generate thrust . A vertical propeller is installed at the nose to provide tension . It has a v-shaped tail which can improve both longitudinal and directional stability. The strake is employed to improve the aerodynamic characteristics of the wing at a large angle of attack, as can been seen in Figure 1. 32 978-1-5386-5836-9/18/$31.00 \u00a92018 IEEE DOI 10.1109/IHMSC.2018.00016 The deflected slipstream to increase the lift is enabled by the double slit flaps which change the direction of the propellers slipstream. The resultant force of thrust, lift and drag is supposed to be vertical to offset part of gravity and thus achieving short takeoff and landing [6], as shown in Figure 2. The shafts of main propellers are parallel with the centerline of the fuselage. In order to get a vertical force for take-off and landing, a certain angle between the body and the ground shall be maintained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002568_s11465-019-0525-2-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002568_s11465-019-0525-2-Figure10-1.png", + "caption": "Fig. 10 Plate theory-based model and potential new models", + "texts": [ + " In a rotor-bearing system dynamic analysis, any \u03b1 and \u03b2 value can be selected to calculate the modulus of elasticity E if only radial or axial stiffness is considered. However, when both stiffnesses are considered, the correct thickness in accordance with the value of \u03b1 \u03b2 must be determined. For an annular plate in the context of this study, the value of \u03b1 \u03b2 must fall within a limited scope. Other models should be considered beyond this limited scope. Generally, \u03b1 \u03b2 > 1 (KaKr) (Fig. 10). Circumstances in which \u03b1 \u03b2 < 1 fall outside the scope of this study, but they will be examined in future research, such as identifying the relevant components in the picture that relate to the scope of this study. This study proposes a new method for the analysis of rotor- bearing system dynamics. The proposed method is based on a novel application of plate theory. The method introduces two new bearing stiffness models (i.e., axial and radial stiffness models) and presents two new concepts, (i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001063_iccas.2008.4694577-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001063_iccas.2008.4694577-Figure1-1.png", + "caption": "Fig. 1 The membership function of subsystem 1.", + "texts": [], + "surrounding_texts": [ + "x\u0307i(t) = Aikxi(t) +Bikui(t) + n\u2211 j=1,j =i Aijkxj(t\u2212 \u03c4ij) yi(t) = Cixi(t) (2) where \u0393k ip is a fuzzy set for p \u2208 In = {1, 2, . . . , q}, k \u2208 Ir = {1, 2, . . . , r}, Aik \u2208 R n\u00d7n and Bik \u2208 R n\u00d7m are the system matrices of the ith subsystem, Aijk is the interconnected matrix with the ith and the jth subsystem. Using the center-average defuzzification, product inference, and singleton fuzzifier, the T\u2013S fuzzy system (2) is inferred as follows: x\u0307i(t) = r\u2211 k=1 \u03bcik(xi(t)) ( Aikxi(t) +Bikui(t) + n\u2211 j=1,j =i Aijkxj(t\u2212 \u03c4ij) ) yi(t) =Cixi(t) (3) where \u03bcik(xi(t)) = \u03c9ik(xi(t))\u2211r k=1 \u03c9ik(xi(t)) \u03c9ik(xi(t)) = q\u220f p=1 \u0393k ip(xip(t)) in which \u0393k ip(xi(t)) is the fuzzy membership grade of xip(t) in \u0393k ip. Hence, a fuzzy system of a decentralized dynamic output feedback controller is described as follows: Controller Rule m: IF xi1 is \u0393m i1 , \u00b7 \u00b7 \u00b7 , and xiq is \u0393m iq , THEN ui(t) = Cc ikx c i (t) xc i (t) = Ac ikx c i (t) +Bc ikyi(t) (4) where Ac ik, Bc ik and Cc ik are the control gain matrices of appropriate dimension. The defuzzified output of the controller (4) is described by ui(t) = r\u2211 k=1 \u03bcik(xi(t))Cc ikx c i (t) xc i (t) = r\u2211 k=1 \u03bcik(xi(t)) ( Ac ikx c i (t) +Bc ikyi(t) ) (5) Substituting (5) into (3), we can obtain the ith closedloop system of the nonlinear interconnected system with the dynamic output feedback controller. This T\u2013S fuzzy system can be rearranged in following form: [ x\u0307i(t) x\u0307c i (t) ] = r\u2211 k=1 r\u2211 m=1 \u03bcik(xi(t))\u03bcim(xi(t)) \u00d7 ( [ Aik BikC c im Bc imCi Ac im ] [ xi(t) xc i (t) ] + n\u2211 j=1,j =i [ Aijk 0 ] xj(t\u2212 \u03c4ij) ) (6) 3. MAIN RESULTS The objective of this paper is to guarantee the stability of the closed-loop system (6) and obtain the control gain matrices Ac ik, Bc ik, Cc ik and Dc ik. In order to obtain the control gain easily, the following lemmas are required. Lemma 1 ([5]) : Given constant symmetric matrices N , O and L of appropriate dimensions, the following inequalities O > 0, N + LTOL < 0 are equivalent to the following inequality[ N LT L \u2212O\u22121 ] < 0 or [\u2212O\u22121 L LT N ] < 0 . Lemma 2 ([11]) : For real matrices X and Y with appropriate dimensions, the following inequality is always satisfied. XTY + Y TX \u2264 \u03c3XTX + 1 \u03c3 Y TY where \u03c3 is a positive constant. For guaranteeing the stability condition, we define Lyapunov functional candidate as follows: V (x(t)) = n\u2211 i=1 Vi(xi(t)) Vi(xi(t)) =xT i (t)Pixi(t) + n\u2211 j=1,j =i \u222b t t\u2212\u03c4ij xT j (s)xj(s)ds (7) where Pi is a matrix with symmetric and positive for i \u2208 In = {1, 2, . . . , n}. Clearly, V (t) is positive and radially unbounded. Form Lyapunov candidate (7), the asymptotic stability condition for the nonlinear interconnected system with time delay is represented in the following theorem. Theorem 1: If there exist a positive definite matrices Xi, Yi and some matrices Ac ikm, Bc im, Cc im, Ni, such that the following LMIs and equations are satisfied, then the whole closed-loop nonlinear interconnected system (6) is asymptotically stable at the its equilibrium point.[ Xi I I Yi ] 0 (8) \u23a1 \u23a3 1 n\u22121 (\u0398ikm + \u0398T ikm) \u2217 \u2217 A\u0303T ijk \u2212I \u2217 Gi 0 \u2212I \u23a4 \u23a6 \u227a 0 (9) for i, j \u2208 In = {1, 2, . . . , n} with i = j and k, m \u2208 Ir = {1, 2, . . . , r} with k < m, where \u0398ikm = [ XiAik + B\u0303c imCi A\u0303c ikm Aik AikYi +BiC\u0303 c im ] A\u0303ijk = [ Aijk YiAijk ] Gi = [ I Yi 0 NT i ] A\u0303c ikm =XiAikYi + B\u0303c imCiYi +XiBikC\u0303 c im +MiA c imN T i B\u0303c im =MiB c im C\u0303c im =Cc imN T i and \u2217 denotes the transposed elements in the symmetric positions. Proof: It is omitted in this paper. Remark 1: The inequalities for the pair (k,m) such that \u03bcik(xi(t)\u00d7\u03bcim(xi(t)) = 0, do not need to be solved to determine global stability. 4. SIMULATION We consider the fuzzy model of the interconnected system which is composed of the following subsystems: x\u0307i(t) = 2\u2211 k=1 \u03bcik(xi(t)) ( Aikxi(t) +Bikui(t) + 2\u2211 j=1,j =i Aijkxj(t\u2212 0.2) ) yi(t) =Cixi(t) where x1 = [ x11 x12 ]T , x2 = [ x21 x22 ]T , the system matrices are represented as follows: A11 = [ 5 0 1 \u22125 ] , A12 = [ 4 1 2 \u22123 ] , A21 = [ 2 1 0 \u22128 ] , A22 = [ 1 1 \u22121 \u22124 ] , B11 = [ 1 1 ] , B12 = [ 2 1 ] , B21 = [ 2 5 ] , B22 = [ 1 4 ] , C1 = [ 1 1 ] , C2 = [ 1 1 ] , A121 = [ 0.2 0.3 0.1 0.2 ] , A122 = [ 0.3 0.3 \u22120.1 0.4 ] , A211 = [ 0.2 0.2 0.1 0.3 ] , A212 = [ 0.2 0.3 0.1 0 ] , and the membership functions are shown in Figs. 1 and 2. We can obtain the decentralized dynamic output feedback control gain matrices by using Theorem1 and solv- ing the corresponding LMIs: Ac 11 = [\u221228.5661 \u221236.7599 \u22122.2258 \u22121.3140 ] , Ac 12 = [\u221256.3349 \u221247.8600 2.5933 2.5665 ] , Ac 21 = [\u221236.0003 \u2212118.3699 \u22125.6639 \u221219.1729 ] , Ac 22 = [\u221220.5916 \u221285.9024 \u22124.1634 \u221211.8006 ] , Bc 11 =Bc 12 = [ 8.2188 \u22121.5640 ] , Bc 21 = Bc 22 = [ 16.4733 0.5505 ] , Cc 11 =Cc 12 = [\u221221.4146 \u221210.5872 ] , Cc 21 =Cc 22 = [\u22123.3031 \u22124.8812 ] The initial conditions of the interconnected system are set to x1(0) = [ 1 \u22121 ]T and x2(0) = [ 2 \u22122 ]T . Then, time responses for each subsystem are shown in Figs. 3 and 4. As shown above, all state variables of each subsystems are converged to zero. Thus, we can know that the proposed decentralized dynamic output feedback controller make the nonlinear interconnected system with time delay satisfy an asymptotic stability. 5. CONCLUSIONS In this paper, the decentralized dynamic output feedback controller has been proposed for the nonlinear interconnected systems with time delay. The controller is designed in the T\u2013S fuzzy system form and sufficient conditions for stabilization of the closed-loop system are designed in the LMI format. The numerical example has been shown to prove the advantage of the developed method. REFERENCES [1] X. G. Yan, J. J. Wang and X. Y. Lu\u030b, and S. Y. Zhang, \u201cDecentralised output feedback robust stabilization for a class of nonlinear interconnected systems with similarity\u201d, IEEE Transactions on Automatic Control, Vol. 43, No. 2, pp. 294-299, 1998. [2] X. G. Yan, C. Edwards, and S. K. Spurgeon, \u201cDecentralised robust sliding mode control for a class of nonlinear interconnected systems by static output feedback\u201c, Automatica, Vol. 40, pp. 613-620, 2004. [3] B. Y. Zhu, Q. L. Zhang and X. F. Zhang, \u201cDecentralized robust guaranteed cost control for uncertain T\u2013 S fuzzy interconnected systems with time delays\u201d, International Journal of Information and Systems Sciences, Vol. 1, No. 1, pp. 73-88, 2005. [4] C. S. Tseng and B. S. Chen, \u201cH\u221e decentralized fuzzy model reference tracking control design for nonlinear interconnected systems\u201d, IEEE Transactions on Fuzzy Systems, Vol. 9, No. 6, pp. 795-809, 2001. [5] F. H. Hsiao, C. W. Chen, Y. W. Liang, S. D. Xu and W. L. Chiang, \u201cT\u2013S fuzzy controllers for nonlinear interconnected systems with multiple time delays\u201d, IEEE Transactions on Circuits and Systems, Vol. 52, No. 9, pp. 1883-1893, 2005. [6] R. J. Wang, \u201cNonlinear decentralized state feedback controller for uncertain fuzzy time-delay interconnected systems\u201d, Fuzzy Sets and Systems, Vol. 151, pp. 191-204, 2005. [7] R. J. Wang, W. W. Lin and W. J. Wang, \u201cStabilizability of linear quadratic state feedback for uncertain fuzzy time-delay systems\u201d, IEEE Trans. Syst. Man, Cybern. Park B, Vol. 34, pp. 1288-1292, 2004. [8] J. U. Park, H. Y. Jung, J. I. Park and S. G. Lee, \u201cDecentralized dynamic output feedback controller design for guaranteed cost stabilization of large-scale discrete-time systems\u201d, Applied Mathematics and Computation, Vol. 156, pp. 307-320, 2004. [9] M. Benyakhlef, \u201cDecentralised nonlinear adaptive fuzzy control for a class of large-scale interconnected systems\u201c, International Journal of Computational Cognition, Vol. 4, No. 2, pp. 14-19, 2006. [10] S. S. Stankovic\u0301, D. M. Stipanovic\u0301, and D. D. S\u030ciljak, \u201cDecentralized dynamic output feedback for robust stabilization of a class of nonlinear interconnected systems\u201c, Automatica, Vol. 43, pp. 861-867, 2007. [11] K. Zhou and P. P. Khargonedkar, \u201cRobust stabilization of linear systems with norm-bounded timevarying uncertainty\u201d, Syst. Contr. Lett., Vol. 10, pp. 17-20, 1988." + ] + }, + { + "image_filename": "designv11_92_0001483_978-3-319-72730-1_17-Figure17.14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001483_978-3-319-72730-1_17-Figure17.14-1.png", + "caption": "Fig. 17.14 Flux vectors for torque control", + "texts": [ + " For the rotor flux, on the one hand, we may then derive: d\u03a8rd dt = \u2212 Rr Lr \u03a8rd + Rr Lm Lr Isd (17.9) d\u03a8rq dt = 0; \u03a8rq = Lr Irq + Lm Isq \u2261 0 (17.10) On the other hand, the stator flux can be written as \u03a8sd = \u03c3Ls Isd + Lm Lr \u03a8rd (17.11) \u03a8sq = \u03c3Ls Isq (17.12) while the torque can be written as T = 3 2 Np \u00b7 Im(\u03a8 \u2217 s \u00b7 I s) = 3 2 Np \u00b7 Im(\u03a8 r \u00b7 I \u2217 r ) (17.13) or, T = \u22123 2 Np\u00b7 Lm Lr \u00b7\u03a8rd \u00b7Isq = \u22123 2 Np\u00b7 Lm \u03c3Ls Lr \u00b7\u03a8rd \u00b7\u03a8sq = 3 2 Np\u00b71 \u2212 \u03c3 \u03c3Lm \u00b7|\u03a8 r |\u00b7|\u03a8 s |\u00b7sin \u03b3 (17.14) \u03b3 is the angle between the fluxes \u03a8 s and \u03a8 r , see Fig. 17.14. This torque equation shows that the torque can be controlled by controlling the stator and rotor flux amplitudes and the angle between both vectors. From Eqs. 17.9 and 17.10, we observe that the rotor flux has a very large time constant Lr/Rr (of the order of seconds for medium power machines). The stator flux, in contrast, does not only comprise the rotor flux (corrected with the winding factor-like ratio Lm/Lr ), but also the leakage flux (which may be varied with a small time constant). The stator flux, and thus its amplitude as well as the angle \u03b3, can be changed very quickly by the supply" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003439_cdc.2018.8619197-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003439_cdc.2018.8619197-Figure1-1.png", + "caption": "Fig. 1. 3DOF Robot Arm Model", + "texts": [ + " Moreover, Section III elaborates on how these nonlinear constraints on the strains in the glass plate can be taken into account in a robust MPC design such that a safe pick and place maneuver can be obtained violating none of the constraints in the presence of uncertainties. In Section IV the corresponding developments are illustrated by simulating a realistic KUKA youBot robot arm. Section V concludes the paper. This section is divided into two parts. Section II-A outlines the dynamic model of a robot arm and Section II-B introduces a model of the glass plate that is carried by the robot. 978-1-5386-1395-5/18/$31.00 \u00a92018 IEEE 2647 We consider a robot arm with three degrees of freedom as sketched in Fig. 1. The robot arm is composed of two links and one gripper carrying a glass plate. In the following, ex = [1 0 0]T , ey = [0 1 0]T , ez = [0 0 1]T denote orthogonal basis vectors, where ez is pointing to the sky. There are three joints, which can be used to place the centers of gravity of the links and gripper at R1(\u03b8) = h1 2 ez, R2(\u03b8) = h1ez + l1 2 S1(\u03b81)S2(\u03b82)ey, and R3(\u03b8) = h1ez + ( l1 + l2 2 ) S1(\u03b81)S2(\u03b82)ey , respectively. Here, h1 and l1 are the lengths of the links, l2 the length of the gripper, and \u03b8 = (\u03b81, \u03b82, \u03b83)T stacks the angles of the joints", + " In order to work out the kinetic energy of the robot, we need to work out the angular velocities of all links, the gripper, and the glass plate. They are given by \u03c91(\u03b8\u0307) = 0 0 \u03b8\u03071 , \u03c92(\u03b8, \u03b8\u0307) = \u03b8\u03072 \u03b8\u03071 sin \u03b82 \u03b8\u03071 cos \u03b82 , and \u03c93(\u03b8, \u03b8\u0307) = \u03c9\u0304(\u03b8, \u03b8\u0307) = \u03b8\u03072 cos \u03b83 \u2212 \u03b8\u03071 sin \u03b83 cos \u03b82 \u03b8\u03073 + \u03b8\u03071 sin \u03b82 \u03b8\u03072 sin \u03b83 \u2212 \u03b8\u03071 cos \u03b83 sin \u03b82 . The moment of inertia of the glass plate is assumed to be given by J\u0304 = m\u0304 12 c21 + c22 0 0 0 c22 + c23 0 0 0 c21 + c23 , where m\u0304 = \u03c1c1c2c3 is the mass of the glass plate, \u03c1 its density, and c1, c2 and c3 the lengths of the edges of the plate as sketched in Figure 1. Now, the kinetic energy Ek of the robot and the glass plate is given by Ek = 1 2 3\u2211 i=1 (mi||R\u0307i(\u03b8)||22 + \u03c9i(\u03b8, \u03b8\u0307) TJi\u03c9i(\u03b8, \u03b8\u0307)) + 1 2 (m\u0304|| \u02d9\u0304R(\u03b8)||22 + \u03c9\u0304(\u03b8, \u03b8\u0307)TJ\u0304 \u03c9\u0304(\u03b8, \u03b8\u0307)) . (1) Here, mi and Ji denote the mass and the moments of inertia of robot links as well as the gripper. Similarly, the potential energy is given by Ep(\u03b8) = m1g h1 2 +m2g(h1 + l1 2 sin \u03b82) +m3g(h1 + l1 sin \u03b82 + l2 2 sin \u03b82) + m\u0304g(h1 + l1 sin \u03b82 + l2 sin \u03b82), where g denotes the gravitational constant. The Lagrangian function is given by L(\u03b8, \u03b8\u0307) = Ek(\u03b8, \u03b8\u0307)\u2212 Ep(\u03b8) , (2) i", + " Here, \u201c \u00d7 \u201d denotes the outer product between two vectors in R3. Notice that the above integrals can be worked out explicitly by using state of the art computer algebra tools, but the corresponding expressions are rather long and therefore not displayed as part of this paper. In order to ensure that the glass plate does not break at the gripper, we need to bound the strain of the external fiber at the rectangular area at the front side of the gripper. If cx\u2032 denotes the length of this front edge of the gripper (see Figure 1), the associated rectangular area and section modulus are given by A1 = cx\u2032c2 and S1 = cx\u2032c 2 2 6 . Thus, we must ensure that \u03c3x(s) = |M(s)TS123(\u03b8)Tex| S1 + |F (s)TS123(\u03b8)Tey| A1 \u2264 \u03c4 , (5) where \u03c4 denotes the tensile strength. Similarly, we must ensure that \u03c3y(s) = |M(s)TS123(\u03b8)Tey| S2 + |F (s)TS123(\u03b8)Tex| A2 \u2264 \u03c4 , (6) with A2 = cy\u2032c2 and S2 = cy\u2032c 2 2 6 (see [3] for details on section modulus and tensile strength). Here, cy\u2032 c1 denotes the length of the other edge of the gripper. Nominal time-optimal maneuvers can be found by solving optimal control problems of the form min s, v, T T s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003660_eptc.2018.8654294-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003660_eptc.2018.8654294-Figure2-1.png", + "caption": "Figure 2: Details of Laser Solder Ball Jetting Process", + "texts": [ + " During the laser solder jetting process, a customized ball size conveying disk will rotate and singulate the solder ball out for the reservoir bond head. The singulated solder ball will be transfer down and held at the capillary tip [7-8]. Solder ball will be reflow by the Nd:YAG laser pulse, which is the diode pumped. The molten solder ball was jetting out by the inert purge of the N2 as shown in Figure 3. It\u2019s then, wetted onto the metal pad surface. The jetted solder ball then, cools down immediately upon impact as shown in Figure 2. Molten laser jetted solder ball is then solidified. The solder ball shaped into a hemispherical bump with a good metallurgical bond with the metal bond pad [9]. PacTech SB2-Jet tool was used for the evaluation with direct eutectic AuSn of 120\u03bcm ball size solder with pitch at 200\u03bcm. Image of the silicon based chips with surface finish of Au, Pd & Pt was used as the test vehicle as shown in Figure 4 below. Solder Bump on Au, Pd, Pt Surface The samples are subjected to high temperature storage of 125\u00b0C for 24hrs, 500hrs and 1000hrs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000280_j.jappmathmech.2010.01.002-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000280_j.jappmathmech.2010.01.002-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " We will choose as the generalized coordinates of this rod the angles \u03d1 (0 < \u03d1 < /2) and (0 \u2264 < 2 ), respectively, between the rod and the direction of the force of gravity and between two vertical planes that pass through the fixed point, one of which is fixed and the other contains the rod. The structure of the Lagrangian in these coordinates indicates the presence of two first integrals: the energy integral and the cyclic integral (Kz is the projection of the angular momentum of the system onto the Oz axis, which is directed along the force of gravity, see Fig. 1). These integrals specify two types of motion of the rod: a) if \u0307(0) = 0, the rod moves in a fixed vertical plane as an ordinary physical pendulum b) if \u0307(0) = 0, then \u0307(t) > 0 at any time, and \u03d1(t) is a periodic function of time. Now consider the behaviour of the rod in the case where a Cardan\u2013Hooke universal joint is used to fix its end (Fig. 1). Suppose the plane of the fixed yoke is vertical and the axis of rotation of the cross piece is horizontal. For simplicity, we will assume that the cross piece is a set of two identical uniform thin rods (sections of a straight line), which are rigidly joined at their midpoints at a right angle, and that the depth a of the movable yoke of the cross piece is equal to 0, so that the end of the rod fastened to this yoke is fixed. In this case, the constraints applied to the rod will be exactly the same as in the case considered previously, and if the mass of the centre cross is equal to zero, there will be no special features in its motion under the action of gravity", + "002 Let J ( J = 4 3 ml2 ) be the moment of inertia of the rod about the axis that passes through its end and is perpendicular to it, and let the moment of inertia of the cross piece about its axis of rotation be equal to J ( 1). It is convenient to choose the angles (the angle of rotation of the cross piece, \u2212 /2 < < /2) and (the angle between the rod and the plane that passes through the fixed point and is perpendicular to the axis of rotation of the cross piece, \u2212 /2 < < /2) as the coordinates in this case (Fig. 1). We write the Lagrangian of the system consisting of the rod and the cross piece: This very simple working model of the system exhibits its remarkable behaviour. If we set the rod at first moves near the vertical plane passing through the initial point. Then the angle of deflection from the vertical plane gradually increases, and the rod begins to rotate about the vertical. The rotation gradually changes to planar oscillations, whose plane differs from the original plane. After a certain number of oscillations, the rod again begins to rotate about the vertical, but now in the opposite direction, and so forth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001400_978-3-319-60846-4_23-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001400_978-3-319-60846-4_23-Figure2-1.png", + "caption": "Fig. 2 Greenwood-Williamson rough surface contact geometry", + "texts": [ + " The Greenwood-Williamson model is a statistical contact model based on the Hertzian theory [6], in which the surface roughness is modelled as a cluster of hemispherical asperities. The GW model has the following assumptions [5]: \u2022 The rough surface is considered isotropic. \u2022 Asperities have a spherical shape near their summits. \u2022 All asperity summits have the same radius R, but their heights vary based on the asperity heights distribution function. \u2022 There is no interaction between asperities. \u2022 Bulk deformation is not considered in the model, only asperity deformation. \u2022 Deformation of the asperities is purely elastic. Figure 2 depicts the geometry of a GW rough surface contact. The two rough surfaces in contact are modelled as one smooth surface and one equivalent rough surface with the Root Mean Squared (RMS) roughness is defined by \u03c3 = \u221a \u03c3 2 1 + \u03c3 2 2 , (3) where \u03c31 and \u03c32 are the roughnesses of surface 1 and 2, respectively. Two reference planes are defined: the mean asperity height plane and the mean surface height plane. The former is usually used in the asperity-based contact models and the latter is more practically obtained experimentally [5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002240_msf.928.209-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002240_msf.928.209-Figure3-1.png", + "caption": "Fig. 3. Deflections of bimorph with different Vol% of filler and thickness", + "texts": [ + " These are used to determine the deformation at different temperature of bimorph actuator. The extra finer mesh size using the free triangulation technique is adopted for meshing as shown in Fig.2. COMSOL simulation is performed on the bimorph thermal actuator to investigate the effects of variation of the material properties and thickness with respect to displacement. The deflections of the bimorph are calculated using Equation.2 and finite element (FE) modelling. The comparative study of bimorph deformation at different thickness and volume percentage of CB is shown in Table 3. Fig. 3 depicts a post processor plot of FE analysis. Deformation of bimorph could be visualized from these plots. It could be observed that the deformation is higher for 15 Vol% of CB. The Fig.4 show that displacements of the BM405 obtained from analytical and FE procedures. Results obtained from both approaches mentioned above agree well for lower volume fraction, but differ for higher volume fractions. Results obtained by the analytical route are higher at higher volume fractions as compared to lower volume fractions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000497_amr.76-78.137-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000497_amr.76-78.137-Figure11-1.png", + "caption": "Fig. 11 Schematic diagram of the stiffness of grinding wheel itself under the grinding operation Grinding force (Table feed rate)", + "texts": [ + " 10 Schematic diagram of maximum grain depth of cut g g a Grinding wheel D Workpiece v V and the wheel deformation measured in the stationary state, the elastic deformation of grinding wheel in the stationary state was from several micron to over ten micron. Therefore, calculated maximum grain depth of cut g can be neglected. From this calculated result, it can be regarded that the number of contacting abrasive grain existing on the grinding wheel surface is constant irrespective of the increase and/or the decrease of the table feed rate. Therefore, it is considered that the contact stiffness of the grinding wheel itself under the grinding operation for same wheel depth of cut is constant as shown Figure 11. On a next stage, the reason why the contact stiffness decreases with the increase of the table feed rate as shown Figure 5 is investigated. In the grinding operation, the residual stock removal is occurred by the elastic deformation of grinding wheel, while the workpiece is removed by the grinding wheel. And, the contact stiffness is calculated as the ratio of the normal grinding force to the residual stock removal. Here, it is considered that the residual stock removal of workpiece differs whether the table feed rate is fast or slow", + " From this figure, it is known that the contact stiffnesses under the grinding operation increase with the increase of the number of grinding pass per unit length of workpiece relating the table feed rate. This result indicates the influence of the residual stock removal of workpiece. And, from Figure 13, since the contact stiffness increases in the direct proportion to grinding pass, it is considered that the stiffness of grinding wheel itself under the grinding operation is constant as shown Figure11, while the elastic deformation of wheel is varied by the grinding force. That is, since the stiffness of one abrasive grains is constant to against the different grinding force as shown Figure 8, and the number of contacting abrasive grain under the grinding operation for same wheel depth of cut doesn\u2019t vary as shown Figure 9, it is regarded that the stiffness of grinding wheel itself under the grinding operation also is constant for the different grinding force. While the elastic deformation of wheel differs from the normal grinding force as shown Figure 11, since the residual stock removal in case of slow table feed rate and/or of many grinding pass per unit length of workpiece is small quantity, it is considered that the equivalent contact stiffness under the grinding operation increases with the decrease of the table feed rate. On the other hand, from Figure 13, it was known that the contact stiffness was in direct proportion to the number of grinding pass per unit length as shown above. So, in the grinding operation under the range of elastic area of grinding wheel, if the contact stiffness can be measured under the condition of only one table feed rate, the contact stiffness for other table feed rate can be calculated by linear relation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001503_978-3-319-70939-0_18-Figure18.15-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001503_978-3-319-70939-0_18-Figure18.15-1.png", + "caption": "Fig. 18.15 A rigid flat punch sliding against a half-plane", + "texts": [ + "85) with solution w1(t) = \u2212u0 [ 1 \u2212 exp(\u2212\u039bt) ] , (18.86) where \u039b = \u03bb\u03b2(1 + \u0131 f \u03b2); \u03bb\u03b2 = \u03bb (1 + ( f \u03b2)2) , (18.87) and \u03bb is defined in (18.81). Using this result, Eq. (18.84) can be written w(x, t) = \u03b1 f |V |p0t \u2212 u0 exp(\u2212\u03bb\u03b2 t) cos(m(x \u2212 ct)), (18.88) where c = f \u03b2\u03bb\u03b2 m = \u03b1 f |V |E\u2217 f \u03b2 2(1 + ( f \u03b2)2) (18.89) is the migration speed. Notice that the direction of migration depends on the sign of the coupling constant \u03b2. Also, the approach to the steady state is slightly slower than in the uncoupled case, since \u03bb\u03b2 <\u03bb. Figure18.15 shows a rigid initially flat punch of width 2a that is pressed against an elastic half-plane by a force P . The half-space moves at speed V causing wear on the punch. We shall assume that the wear of the half-space that occurs during a single pass under the punch is small enough to be neglected. Suppose that there is no coupling (\u03b2=0), in which case the initial contact pressure distribution [before wear occurs] is given by Eq. (6.14) as However,wearwill occur preferentially near the edges x=\u00b1a until after a sufficiently long time, the contact pressure is uniform and equal to P/2a" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003415_phm-chongqing.2018.00022-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003415_phm-chongqing.2018.00022-Figure5-1.png", + "caption": "Fig. 5 Gear tooth spalling in a straight line", + "texts": [ + " For a healthy tooth, xtA , xA and ztI , zI are given by ( )( ) ( )( ) 3 3 2 , 1 12 2 2 , 1 12 2 xt xt x x xt zt x z A A h L I h L h L I h L = = = = , (20) Where xth is the half tooth thickness of the transaction part evaluated by Eq.(2). xh is the half tooth thickness of an arbitrary point on the involute curve given by Eq.(1). Within the faulty area, xA and zI are given by the values evaluated by Eq.(9) and (13), respectively. The TVMS of a gear pair with tooth spall are functions of the cross-section area Axs, the area moment of inertia Izs, and the effective tooth surface contact length Le. The variation of these parameters modifies the magnitudes of the TVMS. Fig. 5 and Fig. 6 shows the modeled gear tooth spall under different conditions, and the corresponding variations of Axs and Izs. For spur gears, tooth spalls usually occurr near the tooth pitch point and stay in line due to the line contact property of spur gears (see Fig. 1). For simplicity, the tooth spalls in this work are assumed identical and in line. The parameters for the tooth spall are xsk_start=31 (mm), wsk=2 (mm), hsk_max=0.9 (mm) and sk=17.48 . The parameters for the gear pair are shown in table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001725_icit.2018.8352179-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001725_icit.2018.8352179-Figure11-1.png", + "caption": "Fig. 11. Experimental setup in case of dynamic situation", + "texts": [ + " The purpose of this experiment is position estimation of the drone, thus, the position of the drone is not controlled by using estimated position from the tensile force of the tube. The drone keeps hovering using AR.drone\u2019s controller after being moved to a certain approximated position by a human operator. In order to obtain the actual position of the drone, a marker was attached to the side of the drone, and the position of the marker is measured using Kinect. The estimated position is verified by comparing with the actual position. The scene of the experiment is shown in Fig. 11. First, the tensile force and angle of the tube are measured using the measuring device when the drone is hovering. At the same time, the position of the drone is measured using Kinect. Then, the force and its angle are estimated using catenary theory. These estimated values are the estimated position of the drone. Here, since the drone is towing the tube, the position of the drone slightly moves during hovering. Thus, the estimated position is affected by the slight move of the drone. Therefore, it is sufficient for the estimated position to exist within the range of the slight move of the drone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003681_978-3-030-04975-1_89-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003681_978-3-030-04975-1_89-Figure4-1.png", + "caption": "Fig. 4. Discrete models of the gear pump bodies a) external gear pump b) internal gear pump c) gerotor pump d) HEXA element", + "texts": [ + " Thrusts pb1, pb2 of the bearings on the central body are determined as: pb1 \u00bc Fs1 2 db l ; pb2 \u00bc Fs2 2 db l where: db \u2013 bearing diameter, l \u2013 bearing length Mechanical loads generated by force Q the screw clamp are determined as: Q \u00bc M 0; 5 ds tg\u00f0c\u00fe q0\u00de where: M \u2013 screw torque, ds \u2013 thread diameter, c \u2013 helix angle, q\u2019 \u2013 apparent friction angle Using [28\u201331], numerical models of the bodies for the three analysed pump types were developed. The views of the bodies with the assumed finite element mesh are presented in Fig. 4a, b, c. For the making of the mesh, the 1st-order HEXA elements of the same size were used (Fig. 4d). A HEXA element is a cubic element. The element features eight nodes and is a 1st-order element. In each node, the element has three degrees of freedom which are shifts ux, uy, uz relative to the coordinate axes. Figures 2 and 4 show that the capacity of the bodies varied. It was the highest in the case of the external gear pump, average for the internal gear pump, and the lowest for the gerotor pump. Respectively, the number of the finite elements in the mesh of the pumps at the same size of the elements decreased, and was; \u2022 559445 for the external gear pump, \u2022 392348 for the internal gear pump, \u2022 315314 for the gerotor pump" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002082_1.4040813-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002082_1.4040813-Figure3-1.png", + "caption": "Fig. 3 Flooded, pressurized housing oil flow path features [11]", + "texts": [ + " Each pad is centrally positioned on a ball-in-socket, spherical pivot where the ball is constructed of steel and the pivot is constructed of brass. The pads are housed in a vintage flooded, pressurized housing design with oil inlet nozzles\u2014the operating principles of which are described in detail by Nicholas [11]. Oil is introduced to the bearing cavity through the oil supply nozzles, located between each of the pads. The majority of the hot sump oil exits the bearing through two smaller discharge holes located on either side of each supply nozzle, while some leaks axially through two tight clearance floating end seals. Figure 3 provides an illustration of the oil flow path features. ISO VG 32 oil was supplied to the bearings at 43.3 C. Flow rate to each bearing housing was controlled through two valves located just outside each housing and monitored via hydraulic oil flowmeter. Thermocouples embedded in the pads and labeled \u201cTC\u201d were not utilized in the current study. A summary of the important test bearing parameters can be found in Table 1. The primary purpose of this study was to investigate the effects of reduced oil supply flow rates to supporting tilting-pad bearings on damped natural frequency and system damping ratio under various operating conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000036_detc2009-87370-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000036_detc2009-87370-Figure10-1.png", + "caption": "Figure 10. Geometry model of a layer for deposition and for machining, respectively", + "texts": [ + " When building the overhang area based on thin wall, the system need do multi-axis simultaneous movement. Rotation reduces the speed greatly. Second, the build of layers needs multi-axis machining to achieve sloped layers. The machining time becomes an important portion in the whole build time. In order to determine the exact build time, the method to calculate deposition time and machining time are studied respectively. The layers in multi-axis LAMP are called 3D layers because the layers are unnecessary to have uniform thickness. The layer geometry for deposition is shown in Fig. 10(a). Currently it is assumed that the deposition traverse speed for the overlap area, overlapv , and the deposition traverse speed for the overhang area built in thin wall, overhangv , are constant for a specific material. However this is not true for actual deposition. Many factors are involved in the selection of the traverse speed. Deposition for overlap area only needs the translation, which can be very fast. But deposition for overhang area in thin wall needs multi-axis simultaneous movement, especially the rotation, which is much slower than the translation", + " This speeds up the deposition process by making very thick layers during the LM phase of the integrated procedure. After this, the tool-paths generated using a cuspheight, 1, much lower than the used for calculating the thickness of the LM layers. This increases the machining time, but the surface accuracy can be improved. The tool-paths for machining are multiple. The tool-paths enable the manufacture of an object to accuracy greater than what can be achieved by using the deposition process alone. The multi-pass milling tool-paths are shown in Fig. 10(b). Two equations have been used for calculating the machining time. One equation used for calculating the real cutting time is based on length of tool path and the cutting feed rate [19]: C T R V L T (26) where TR is the real machining time, LT is the length of tool path and VC is the cutting feed rate and is constant. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/idetc/cie2009/70870/ on 04/11/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use 8 Copyright \u00a9 2009 by ASME The other equation used for calculating the tool moving time between tool paths is based on the length of non-cutting traverse of cutting tool and the positioning rapid traverse rate: P B M V L T (27) where TM is the traverse time, LB is the length of non-cutting traverse, and VP is the positioning rapid traverse rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000501_fie.2009.5350615-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000501_fie.2009.5350615-Figure8-1.png", + "caption": "FIGURE 8 DEVICE FOR PREGN RA LEONE), (COU", + "texts": [ + " The stud ith the types line were ask reciation for t input to the cted to choose nce a decision n the locati situation, in ue (Figure 6) uring war (Fig nd pediatrics ( clinics response r for seven y two weeks the two facul strial designer ted their desig weeks, each te design to the entatives from mation about he location an roup had to d and how the d ease of disin und other desig In previous gn problem, of the design functional o gled with how rest and still le Ultrasound alth care situa l diversity in t E/IEEE Fron two approach ifferent way iven much m elop one conc gether with ite. The stude s (a Plast ndustrial Des ct. The stude f questions, a noSite\u2019s facil d a tour of o get a bet better understa ound technolo students\u2019 caro ents commen of questions t ing and that th he knowledge project from a location an was made, ea on\u2019s economi frastructure, e ure 7) Figure 8) weeks on th , designers a ty members (o ) would prov n. am made a fi entire class, SonoSite. T how the dev d condition th explain how assembly wo fection. n projects w projects wh by the time s had a simi bjectives of to make th meet the des project, hav tion in a rem he final design tiers in Educ M3E-5 es, to ore ept the nts ics ign nts nd ity, the ter nd gy tid ted hat ey of the d a ch cs, tc. eir nd ne ide nal the he ice ey the uld ere ere the lar the eir ign ing ote s. PORTA PORTAB PO Oc ation Confere BLE ULTRASOUND (HIMALAYAS E ULTRASOUND D (DARFUR, SUDAN LE ULTRASOUND OR REGIONS (SIER tober 18 - 21, nce 2009, San An D FOR HIGH ELEVA SEN, SPOFFORD) A TRIAGE TENT DU CLAIRE, MCCORM ANCY AND PEDIA RTESY WILSON, B tonio, TX TION USE RING ACK) TRIC USE IN LAIR) 978-1-4244-4714-5/09/$25" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000034_20080408-3-ie-4914.00002-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000034_20080408-3-ie-4914.00002-Figure3-1.png", + "caption": "Fig. 3. Six degrees of freedom of motion of a rigid body", + "texts": [ + " A series of trials to identify the dynamic model of the vehicle for controller design evidently proved that the umbilical was causing significant amount of drag and extra weight on the AUV (Naeem, 2004). Furthermore, the umbilical limits the operation of the vehicle since a surface platform was required to move along with the vehicle because of the finite cable length. Hence the need to replace the cable is of paramount importance to test the autonomy of the whole system. Nevertheless, the system identification trials were carried out with the umbilical with care exercised to minimise its effects on the data. An analysis of the captured data is now presented. Fig. 3 depicts the six degrees of freedom of motion of an underwater vehicle where O is the origin or centre of gravity of the AUV. The variables of interest include yaw, pitch, roll and the linear velocities along the three axes. In this paper, since the main focus is on the lateral motion of the vehicle therefore the depth parameter is not considered. It is however vital to analyse any cross-coupling effect between the variables given in Fig. 3. Fig. 4 shows one of the data sets obtained from Hammerhead through trials conducted at Roadford Reservoir in Devon, UK. The control input or rudder demand is in the form of a pseudo-random binary sequence (PRBS) whereas the observed output is the heading or yaw of the vehicle showing clear patterns of input and output relationship. The pitch and roll variables are also presented and analysed in Fig. 5. It is clear that there is almost no change in pitch and roll with respect to the change in the heading angle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003877_ijthi.2019070103-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003877_ijthi.2019070103-Figure6-1.png", + "caption": "Figure 6. Avoidance with varying velocity: (a) collision at t1 ( )V V a b", + "texts": [], + "surrounding_texts": [ + "In\ufeffthe\ufefftheory\ufeffof\ufeffcomputer\ufeffsimulation,\ufeffthe\ufeffpositions\ufeffof\ufeffvirtual\ufeffmembers\ufeffare\ufeffconfirmed\ufeffby\ufeffcoordinates.\ufeff A\ufefftwo-dimensional\ufeffplane\ufeffis\ufeffconstructed\ufeffto\ufeffsimulate\ufeffthe\ufeffincident\ufeffsite\ufeffin\ufeffthis\ufeffpaper.\ufeffEvery\ufeffposition\ufeffon\ufeff the\ufeffplane\ufeffhas\ufeffa\ufeffspecific\ufeffcoordinate\ufeff ( , )x y ,\ufeffcorresponding\ufeffto\ufeffeach\ufefflocus\ufeffof\ufeffevery\ufeffvirtual\ufeffplayer.\ufeffThe\ufeff coordinate\ufeffvariety\ufeffindicates\ufeffchanging\ufeffof\ufeffthe\ufeffplayer\u2019s\ufeffposition,\ufeffand\ufeffthe\ufeffdifference\ufeffof\ufeffthe\ufeffcoordinate\ufeff values\ufeffrepresents\ufeffthe\ufeffdistance\ufeffbetween\ufefftwo\ufeffplayers.\ufeffWhen\ufeffthe\ufeffformation\ufeffchanges,\ufeffthe\ufeffnext\ufeffstep\ufeffforward\ufeff of\ufeffthe\ufeffmembers\ufeffcan\ufeffbe\ufeffobtained\ufeffon\ufeffthe\ufeffbasis\ufeffof\ufeffthe\ufefforiginal\ufeffand\ufeffgoal\ufeffcoordinates. This\ufeffpaper\ufeffregards\ufeffthe\ufefflocus\ufeffwhere\ufeffthe\ufeffgroup\ufeffformation\ufefflocates\ufeffas\ufeffa\ufefftwo-dimensional\ufeffplane.\ufeffThe\ufeff moving\ufeffdirections\ufeffof\ufeffthe\ufeffvirtual\ufeffplayers\ufeffare\ufeffindicated\ufeffby\ufeffthe\ufeffincrements\ufeff \u2206x \ufeffand\ufeff \u2206y \ufeffin\ufeffx-axis\ufeffand\ufeff y-axis\ufeff coordinate\ufeff directions.\ufeff The\ufeff values\ufeff of\ufeff \u2206x \ufeff and\ufeff \u2206y \ufeff are\ufeff continuously\ufeff recorded\ufeff at\ufeff a\ufeff certain\ufeff frequency,\ufeffbase\ufeffon\ufeffwhich\ufeffthe\ufeffpath\ufeffof\ufeffthe\ufeffvirtual\ufeffplayers\ufeffcould\ufeffbe\ufeffknown.\ufeffUsing\ufeffthe\ufeffpath\ufeffconversion\ufeff algorithm,\ufeffthe\ufeffentire\ufeffconversion\ufeffprocedure\ufeffcould\ufeffbe\ufeffsimulated\ufeffto\ufeffdisplay\ufeffthe\ufeffdetailed\ufefftrajectory\ufeffof\ufeff every\ufeffplayer\u2019s\ufeffpath\ufeffchange. Every\ufeffgroup\ufeffis\ufeffrequired\ufeffto\ufeffkeep\ufeffaggregated\ufeffand\ufeffsustainable\ufeffafter\ufefftransforming\ufeffformation.\ufeffThereby,\ufeff the\ufefftransformed\ufeffformation\ufeffis\ufeffdivided\ufeffinto\ufeffpolygon\ufeffregions\ufeffbased\ufeffon\ufeffthe\ufeffquantity\ufeffof\ufeffleaders.\ufeffFirst,\ufeff the\ufeffclosest\ufeffpoint\ufeffto\ufeffthe\ufeffregional\ufeffcenter\ufeffis\ufeffset\ufeffas\ufeffthe\ufeffposition\ufeffof\ufeffleader,\ufeffand\ufeffthen,\ufeffbased\ufeffon\ufeffgreedy\ufeff algorithm,\ufeffleader\ufeffare\ufeffassigned\ufeffto\ufeffdifferent\ufeffregions\ufeffto\ufefffeedback\ufeffreal-time\ufefflocation,\ufefffinally,\ufeffcollision\ufeff will\ufeffbe\ufeffpredicted,\ufeffdetected\ufeffand\ufeffavoided\ufeffin\ufeffthe\ufeffmovement. Formation control with Greedy Algorithm Based on conditional Formation Feedback Aiming\ufeffat\ufeffcoordinating\ufeffthe\ufeffactions\ufeffbetween\ufeffobstacle\ufeffavoidance\ufeffand\ufeffformation\ufeffkeeping,\ufeffwhile\ufeffone\ufeff player\ufeff is\ufeff avoiding\ufeff obstacles,\ufeff the\ufeff others\ufeff must\ufeff keep\ufeff the\ufeff formation\ufeff operating\ufeff orderly\ufeff based\ufeff on\ufeff the\ufeff greedy\ufeffalgorithm\ufeff(Pan\ufeff&\ufeffRuiz,\ufeff2014).\ufeffMembers\ufeffdon\u2019t\ufeffhave\ufeffany\ufeffrelations\ufeffwhile\ufeffmaintaining\ufeffformation\ufeff because\ufeff of\ufeff the\ufeff parallel\ufeff structure,\ufeff which\ufeff makes\ufeff formation\ufeff control\ufeff more\ufeff flexible\ufeff and\ufeff stable\ufeff (Xin,\ufeff Daqi\ufeff&\ufeffYangyang,\ufeff2017).\ufeffIn\ufeffthe\ufeffmovement,\ufeffFormation\ufefffeedback\ufeffis\ufeffconducted\ufeffin\ufefforder\ufeffto\ufeffresolve\ufeffthe\ufeff issue\ufeffthat\ufeffsome\ufefffollower\ufeffmay\ufeffbe\ufeffout\ufeffof\ufeffthe\ufeffrange\ufeffof\ufeffcommunication\ufeffbecause\ufeffof\ufeffobstacle\ufeffavoidance.\ufeff Consequently,\ufeffit\ufeffprobably\ufeffleads\ufeffto\ufeffthe\ufeffobstruction\ufeffin\ufeffthe\ufeffwhole\ufeffmovement,\ufeffand\ufeffsystem\ufeffwill\ufeffstay\ufeffthe\ufeff locked\ufeffstate\ufeffwhile\ufeffmultiple\ufefffollowers\ufeffare\ufeffstuck\ufeffin\ufeffbarrier\ufeffsimultaneously.\ufeffThereby,\ufeffthe\ufeffintroduction\ufeff of\ufeffconditional\ufefffeedback\ufeff(Singh\ufeff&\ufeffAhmed,\ufeff2014)\ufeffis\ufeffto\ufeffsolve\ufeffthis\ufeffproblem,\ufeffas\ufeffshown\ufeffin\ufeffFigure\ufeff3. Simulate\ufeffa\ufefftwo-dimensional\ufeffplane\ufeffwith\ufeffn\ufeffpoints\ufeffrepresenting\ufeffthe\ufeffleaders,\ufeffand\ufeffimport\ufeffa\ufeffmatrix\ufeff whose\ufeffitems\ufeffrepresent\ufeffthe\ufeffdistances\ufeffbetween\ufeffthe\ufefforiginal\ufefflocation\ufeffand\ufeffthe\ufeffgoal\ufeffof\ufeffleaders.\ufeffIn\ufeffthis\ufeffway,\ufeff the\ufeffproblem\ufeffof\ufeffcalculating\ufeffthe\ufeffshortest\ufeffroute\ufeffwill\ufeffbe\ufefftransmuted\ufeffinto\ufeffan\ufeffissue\ufeffof\ufeffsearching\ufeffthe\ufeffminimum\ufeff value\ufeffin\ufeffthe\ufeffmatrix,\ufeffwhich\ufeffis\ufeffmarked\ufeffLn n a ij\u00d7 = ( ) .\ufeffWhere\ufeff i \ufeffrepresents\ufeffthe\ufeffnth\ufeffleader\ufeffof\ufefforiginal\ufeff position;\ufeff j \ufeffstands\ufefffor\ufeffthe\ufefflocation\ufeffof\ufeffthe\ufeffi-th\ufeffleader\ufeffin\ufeffthe\ufeffobject\ufeffformation;\ufeffa ij \ufeffis\ufeffthe\ufeffdistance\ufefffrom\ufeff the\ufefforiginal\ufefflocation\ufeffto\ufeffthe\ufeffgoal\ufeffof\ufeffthe\ufeffi-th\ufeffleader.\ufeffWhile\ufefftransformation,\ufeffall\ufeffthe\ufeffvirtual\ufeffplayers\ufeffmove\ufeff orderly\ufeffand\ufeffthen\ufeffthe\ufeffwhole\ufeffformation\ufeffcould\ufeffbe\ufefftransformed\ufefffrom\ufeffinitial\ufeffshape\ufeffto\ufefftarget\ufeffshape.\ufeffExcept\ufeff Algorithm 1. Example of a greedy algorithm Void\ufeffContainerLoading(int\ufeffx[],float\ufeffw[],float\ufeffc,int\ufeffn)\ufeff {\ufeff//x[i]=1\ufeffWhen\ufeffand\ufeffonly\ufeffwhen\ufeffthe\ufeffcontainer-i\ufeffis\ufeffloaded\ufeffon\ufeffthe\ufeffweight,\ufeffsorted\ufeffby\ufeffindirect\ufeffaddressing\ufeffmode\ufeff int*t=new\ufeffint[n+1];\ufeff//table\ufeffn\ufeffis\ufeffindirect\ufeffaddressing\ufeff IndirectSort(w,t,n);\ufeff//now,\ufeffw[t(i)]\ufeffw[t(i+1)],1\ufeffi\ufeffn\ufeff \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0for(int\ufeffi=1;i<=n;i++)//initialize\ufeffx\ufeff \u00a0\u00a0\u00a0\u00a0\u00a0\u00a0X[i]=0;\ufeff for(i=1;i<=n)\ufeff&&\ufeffw[t(i)]<=c;i++{\ufeff//select\ufeffitems\ufeffby\ufeffweight\ufeff x[t(i)]=1;\ufeff c=w[t(i)];\ufeff }\ufeff//residual\ufeffcapacity\ufeff delete\ufefft[]; for\ufeffthe\ufeffmovement\ufeffroutes,\ufeffThe\ufeffposition\ufeffof\ufeffeach\ufeffmember\ufeffis\ufefffixed\ufeffboth\ufeffin\ufefforiginal\ufeffand\ufeffgoal\ufeffformation,\ufeff suppose\ufeffthe\ufeffdistance\ufeffis\ufeffa\ufeffroute\ufeffbetween\ufeffthe\ufeffstarting\ufeffpoint\ufeffand\ufeffthe\ufefffoothold,\ufeffand\ufeffthe\ufeffpriority\ufeffalgorithm\ufeff aims\ufeffto\ufefflook\ufefffor\ufeffan\ufeffideal\ufeffroute\ufeffwhich\ufeffis\ufeffthe\ufeffshortest\ufeffone\ufeffwithout\ufeffany\ufeffconflict\ufeff(Bj\u00f8rn\u00f8y,\ufeffBassett\ufeff&\ufeff Ucar,\ufeff2016;\ufeffMa,\ufeffTang\ufeff&\ufeffWang,\ufeff2016).\ufeffA\ufeffmember\ufeffcan\ufeffbe\ufeffaccomplished\ufeffthe\ufefftask\ufeffby\ufeffrunning\ufeffforward\ufeff from\ufeffthe\ufeffinitial\ufeffpoint\ufeff to\ufeffthe\ufeffgoal\ufeffformation.\ufeffAccording\ufeffto\ufeffthe\ufeffdistance\ufeffcalculated\ufefffrom\ufeffthe\ufeffinitial\ufeff location\ufeffto\ufeffthe\ufeffgoal,\ufeffa\ufeffdistance\ufeffmatrix\ufeffof\ufeffn n\u00d7 \ufeffis\ufeffobtained. 26 15 37 20 12 11 14 21 19 12 9 13 8 24 11 23 \u2192 26 15 11 37 20 21 9 13 23 \ufeff (6) So\ufeffthe\ufeffchange\ufeffin\ufeffthe\ufeffthis\ufeffmatrix\ufeffindicates\ufeffthat\ufeffthe\ufeffshortest\ufeffroute\ufeffis\ufeffa 33 8= . collision Prediction The\ufeffpositions\ufeffof\ufeffthe\ufeffwhole\ufeffformation\ufeffmembers\ufeffin\ufeffthe\ufeffgoal\ufeffformation\ufeffare\ufeffcalculated\ufeffusing\ufeffconditional\ufeff feedback\ufeff(Andrade\ufeff&\ufeffSantos,\ufeff2017;\ufeffBarron,\ufeffCohen\ufeff&\ufeffDahmen,\ufeff2008).\ufeffHowever,\ufeffin\ufeffthe\ufeffwhole\ufeffprocess\ufeff of\ufeffthe\ufeffformation\ufefftransformation,\ufeffit\ufeffis\ufeffinevitable\ufeffthat\ufefftwo\ufeffor\ufeffmore\ufeffmembers\ufeffcome\ufeffacross\ufeffsimultaneously,\ufeff and\ufeffit\ufeffis\ufeffcollision.\ufeffIf\ufeffthe\ufeffproblem\ufeffhad\ufeffnot\ufeffbeen\ufeffhandled\ufeffproperly,\ufeffthe\ufeffwhole\ufeffformation\ufeffwould\ufeffhave\ufeff be\ufeffat\ufeffa\ufeffstandstill\ufeff(Niewiadomska-Szynkiewicz\ufeff&\ufeffSikora,\ufeff2011).\ufeffAccordingly,\ufeffcollision\ufeffneeds\ufeffto\ufeffbe\ufeff predicted\ufeffthrough\ufeffthe\ufeffwhole\ufeffprocess. In\ufeffthis\ufeffstudy,\ufeffx \ufeffstands\ufefffor\ufeffa\ufeffformation\ufeffmember,\ufeffL x( ) stands\ufefffor\ufeffthe\ufeffpresent\ufefflocation\ufeffof\ufeffx ,\ufeffV x( ) \ufeff is\ufeffthe\ufeffvelocity\ufeffof\ufeff x ,\ufeffand\ufeffW x( ) \ufeffis\ufeffthe\ufeffwidth\ufeffof\ufeffregion,\ufeffwhich\ufeffis\ufeffa\ufeffconstant\ufeffin\ufeffthis\ufeffstudy\ufeff.\ufeffSuppose\ufeff that\ufeff2\ufeffadjacent\ufeffplayers\ufeffare\ufeffa \ufeffand\ufeffb ,\ufeffand\ufeffL L a L b r = ( ) ( )- ,\ufeffV V a V b r = ( ) ( )- ,\ufeffwhere\ufeffL r \ufeffstands\ufefffor\ufeffthe\ufeff relative\ufeffposition\ufeffof\ufeffa \ufeffand\ufeffb ,\ufeffand\ufeffV r \ufeffthe\ufeffrelative\ufeffvelocity\ufeffof\ufeffa \ufeffand\ufeffb ,\ufeffso\ufeffa\ufeffprobable\ufeffcollision\ufeffbetween\ufeff a \ufeffand\ufeffb \ufeffshould\ufeffsatisfy: L L V t V t W a W b r r r r 2 2 2 22+ \u00d7 \u00d7 \u00d7 + \u00d7 = + +( )( ) ( ) \u03b5 \ufeff (7) Where\ufeff\u03b5 \ufeffis\ufeffthe\ufeffsafe\ufeffdistance\ufeffof\ufeffa \ufeffand\ufeffb .\ufeffThe\ufeffexistence\ufeffof\ufeffunique\ufeffsolution\ufeffor\ufeffunbounded\ufeffsolution\ufeff predicts\ufeffthat\ufeffa\ufeffwould\ufeffnot\ufeffcollide\ufeffwith\ufeffb.\ufeffIf\ufefftwo\ufeffsolutions\ufeffof\ufefft 1 \ufeffand\ufefft 2 t t 1 2 <( ) \ufeffexist,\ufeffit\ufeffindicates\ufeffthat\ufeff there\ufeffwill\ufeffbe\ufeffa\ufeffcollision\ufeffinstantly\ufeff(so\ufeffcollision\ufeffavoidance\ufeffis\ufeffessential\ufeffin\ufeffthis\ufeffcase)\ufeff;\ufeffif\ufefft 2 0< ,\ufeffa\ufeffwould\ufeff collide\ufeff with\ufeff b\ufeff after\ufeff t 1 .\ufeff The\ufeff time\ufeff of\ufeff collision\ufeff could\ufeff be\ufeff indicated\ufeff uniformly\ufeff with\ufeff t p ,\ufeff then\ufeff L L a V a t a p = + \u00d7( ) ( ) ,\ufeffL L b V a t b p = + \u00d7( ) ( ) ,\ufeffwhere\ufeffL a \ufeffand\ufeffL b \ufeffstands\ufefffor\ufeffthe\ufeffpositions\ufeffof\ufeffa \ufeffand\ufeff b \ufeffafter\ufeffthe\ufefftime\ufeff t p \ufeffrespectively,\ufeffand\ufeffthen: 1.\ufeff\ufeff L L V a a b \u2212( )\u00d7 <( ) 0 \ufeffrepresents\ufeffrear\ufeffcollision, 2.\ufeff\ufeff L L V a a b \u2212( )\u00d7 >( ) 0 \ufeffand\ufeffV a V b( ) ( )\u00d7 < 0 \ufeffrepresent\ufefffront\ufeffcollision, 3.\ufeff\ufeff L L V a a b \u2212( )\u00d7 >( ) 0 \ufeffand\ufeffV a V b( ) ( )\u00d7 \u2265 0 \ufeffrepresent\ufeffrear\ufeffcollision, 4.\ufeff\ufeff V b( ) = 0 \ufeffrepresents\ufeffstationary\ufeffcollision. The\ufefflocalized\ufeffcollision\ufeffavoidance\ufeffalgorithm\ufeffis\ufeffdeveloped\ufeffbased\ufeffon\ufeffthese\ufefffour\ufefftypes. collision Avoidance Collision\ufeffavoidance\ufeff(Ajorlou,\ufeffAsadi\ufeff&\ufeffAghdam,\ufeff2015;\ufeffLozano-P\u00e9rez\ufeff&\ufeffWesley,\ufeff1979)\ufeffin\ufeffthis\ufeffpaper\ufeff is\ufeffprimarily\ufeffconducted\ufeffby\ufeffthe\ufeffchange\ufeffof\ufeffvelocity\u2019s\ufeffmagnitude\ufeffand\ufeffdirection.\ufeffSo\ufeffthe\ufeffavoidance\ufeffcan\ufeffbe\ufeff achieved\ufeffwith\ufeffthe\ufeffchange\ufeffof\ufeffmagnitude\ufefflike\ufeffacceleration\ufeffor\ufeffdeceleration\ufeffand\ufeffdirection\ufefflike\ufeffleft\ufeffor\ufeff right,\ufeffsimultaneously\ufeffor\ufeffnot.\ufeffThe\ufeffacceleration\ufeffof\ufeffa \ufeffis\ufeffequal\ufeffto\ufeffthe\ufeffdeceleration\ufeffof\ufeffb ,\ufeffso\ufeffthe\ufeffdeceleration\ufeff rule\ufeffcould\ufeffbe\ufeffrealized\ufeffby\ufeffthe\ufeffacceleration\u2019s.\ufeffThe\ufeffdiagrams\ufeffof\ufeffthe\ufeffleft,\ufeffthe\ufeffright\ufeffand\ufeffthe\ufeffvelocity-vary\ufeff avoidance\ufeffare\ufeffrespectively\ufeffdisplayed\ufeffin\ufeffFigure\ufeff4,\ufeffFigure\ufeff5\ufeffand\ufeffFigure\ufeff6. Formation Initial simulation The\ufeffsimulation\ufeffis\ufeffimplemented\ufeffcombining\ufeffthe\ufeffgreedy\ufeffalgorithm\ufeffand\ufeffconditional\ufefffeedback. 1.\ufeff\ufeff Simulation Environment:\ufeffLength\ufeffl\ufeff=\ufeff200,\ufeffwidth\ufeffw\ufeff=\ufeff200,\ufeffgrid\ufeffsize\ufeffc\ufeff=\ufeff1;\ufeffgrid\ufeffdata\ufeffX\ufeffaxis\ufeff direction\ufeff N x =1/ e ;\ufeff grid\ufeff data\ufeff Y\ufeff axis\ufeff direction\ufeff N w c y = / ;\ufeff X\ufeff axis\ufeff direction\ufeff target\ufeff value\ufeff des N x _ / )x=round( 2 10\u2212 ;\ufeffY\ufeffaxis\ufeffdirection\ufefftarget\ufeffvalue\ufeffdes y _y=N \u221220 . 2.\ufeff\ufeff Two-dimensional\ufeffnetwork\ufeffenvironment\ufeffconstruction:Environment\ufefffunction\ufeffA initialize N N x y = ( , ) \ufeff and\ufeff A ones N N x y = \u00d72 ( , ) .\ufeffThe\ufefffunction\ufeffsets\ufeffthe\ufeffrandom\ufeffposition\ufeffand\ufefftarget\ufeffposition\ufeffof\ufeffthe\ufeff initial\ufeffteam.\ufeffFor\ufeffexample,\ufeffthe\ufefffour\ufeffplayers\u2019\ufeffpositions\ufeffin\ufeffthe\ufefffollowing\ufeffsimulation\ufeffdiagram\ufeffare\ufeff marked\ufeffas\ufefffollowed:\ufeffA(90,50)\ufeff=\ufeff1;\ufeffA(30,30)\ufeff=\ufeff1;\ufeff(50,40)\ufeff=\ufeff1;\ufeffA(100,20)\ufeff=\ufeff1;\ufeffwhere\ufeffthe\ufeffred\ufeff pattern\ufeffrepresents\ufeffthe\ufeffmembers. 3.\ufeff\ufeff The Determination of the Leader:\ufeffFirstly,\ufeffthe\ufeffmembers\ufeffin\ufeffthe\ufeffenvironment\ufeffare\ufeffsearched,\ufeffand\ufeff the\ufeffcoordinates\ufeffand\ufefftheir\ufeffnumbers\ufeffare\ufeffwritten\ufeffdown,\ufeffand\ufeffthen\ufeffthe\ufeffdistance\ufeffbetween\ufeffthe\ufeffmember\ufeff first\ufeffsearched\ufeffand\ufeffthe\ufefftarget\ufeffpoint\ufeff(task\ufefffinal\ufefftarget\ufeffpoint)\ufeffis\ufeffcalculated\ufeffas\ufeffa\ufeffbasis\ufefffor\ufeffcomparison.\ufeff Secondly,\ufeffthe\ufeffdistance\ufeffbetween\ufeffthe\ufefftarget\ufeffpoint\ufeffand\ufeffother\ufeffmembers\ufeffis\ufeffcalculated,\ufeffand\ufeffthen\ufeffthe\ufeff minimum\ufeffdistance\ufeffvalue\ufeffis\ufeffobtained.\ufeffFinally,\ufeffthe\ufeffnearest\ufeffmember\ufefffrom\ufeffthe\ufefftarget\ufeffpoint\ufeffis\ufeffset\ufeffas\ufeff the\ufeffleader. 4.\ufeff\ufeff The Formation of Team Members:\ufeffTo\ufefftake\ufeffa\ufeffdiamond\ufefffor\ufeffexample,\ufefffirst\ufeffdetermine\ufeffall\ufeff the\ufeff geometric\ufeffform\ufeffposition,\ufeffand\ufeffthe\ufeffpoint\ufeffcode\ufeffis\ufeffas\ufeffbelow: position_x\ufeff=\ufeff[leader_xleader_x-40leader_x-40leader_x-80];\ufeffposition_y\ufeff=\ufeff[leader_yleader_y10leader_y+10leader_y].\ufeff Determine\ufeffthe\ufeffgeometric\ufefftarget\ufeffposition\ufeffof\ufeffthe\ufefffour\ufeffmembers.\ufeffFirstly\ufeffthe\ufeffdistance\ufefffrom\ufeffthe\ufeffi-th\ufeff follower\ufeffto\ufeffthe\ufeffleader\ufeffis\ufeffcalculated,\ufeffsecondly\ufefffrom\ufeffthe\ufeffi-th\ufefffollower\ufeffto\ufeffits\ufeffgoal\ufeffpoint,\ufeffultimately\ufeffthe\ufeff closest\ufeffposition\ufeffapart\ufefffrom\ufeffthe\ufeffi-th\ufefffollower\ufeffis\ufeffobtained,\ufeffand\ufeffits\ufeffobject\ufeffposition\ufeffin\ufeffthe\ufeffdiamond\ufeffis\ufeff found.\ufeffAnd\ufeffthe\ufeffcode\ufeffis\ufeffas\ufeffbelow: < (b) a arrives the collision position far before for i=1:N if (r_x(i)==leader_x)&&(r_y(i)==leader_y) r_px(i)=leader_x; r_py(i)=leader_y; position_x(i)=1000; leader=i; end end for i=1:N if i==leader else r_px(i)=position_x(l); r_py(i)=position_y(l); r_p(i)=sqrt((r_x(i)-position_x(l))^2+(r_y(i)position_y(l))^2); for j=2:length(position_x) l=sqrt((r_x(i)-position_x(j))^2+(r_y(i)-position_y(j))^2); if l \u2264( )4 3 4&& ,\ufeffthen\ufeffthe\ufeffmovement\ufeffprocess\ufeffis\ufeffas\ufefffollows:\ufeffthe\ufefffollower\ufeffmoves\ufeffone\ufeff grid\ufeffhorizontally\ufeffwhen\ufeffA r x r y_ , _+( ) ==1 1 ,\ufeffotherwise\ufeffmoves\ufeffone\ufeffgrid\ufeffvertically. grid\ufeffhorizontally\ufeffwhen\ufeffA r x r y_ , _+( ) ==1 1 ,\ufeffotherwise\ufeffretreats\ufeffone\ufeffgrid\ufeffvertically. Figure\ufeff7\ufeffindicates\ufeffthat\ufeffthe\ufeffformation\ufeffcan\ufeffbe\ufeffwell\ufeffmaintained\ufeffintroducing\ufeffformation\ufeffcontrol\ufeffbased\ufeff on\ufeffgreedy\ufeffalgorithm\ufeffwith\ufeffconditional\ufefffeedback\ufeffwhen\ufeffavoiding\ufeffobstacles. sIMULATIoN eXPerIMeNT ANd resULT ANALysIs dynamic Model Based on Leader-Follower Algorithm According\ufeffto\ufeffspecific\ufeffmission\ufeffrequirements,\ufeffunits\ufeffin\ufeffemergencies\ufeffprocess\ufeffcan\ufeffbe\ufeffable\ufeffto\ufeffchange\ufeffand\ufeff maintain\ufeffspecific\ufeffgeometric\ufeffshapes,\ufeffand\ufeffrealize\ufeffformation\ufeffand\ufefftransformation\ufeffin\ufeffreal\ufefftime\ufeffaccording\ufeff to\ufeffchanges\ufeffin\ufeffthe\ufeffenvironment\ufeffto\ufeffbetter\ufeffcomplete\ufeffthe\ufeffassignment\ufeffPeng,\ufeff(Dimarogonas,\ufeffTsiotras\ufeff&\ufeff Kyriakopoulos,\ufeff2009;\ufeffWen\ufeff&\ufeffRahmani,\ufeff2013).\ufeffFor\ufeffpractical\ufeffproblems,\ufeffthe\ufeffformation\ufeffcan\ufeffbe\ufeffvaried\ufeff according\ufeffto\ufeffthe\ufeffenvironmental\ufefffactors\ufeffand\ufeffthe\ufeffdifficulty\ufeffof\ufeffdifferent\ufefftasks.\ufeffIn\ufeffdealing\ufeffwith\ufeffemergencies,\ufeff the\ufeffformation\ufeffof\ufeffthe\ufefftransformation\ufeffoften\ufeffuse\ufeffa\ufeffdiamond-shaped\ufeffdeformation\ufeffH-shaped\ufeffformation,\ufeff horizontal\ufeff team\ufeffformation\ufeffdiamond-shaped\ufeffformation,\ufeffdiamond-shaped\ufeffteam\ufeffdeformation\ufeffand\ufeffso\ufeff on(Lin,\ufeffHwang\ufeff&\ufeffWang,\ufeff2014;\ufeffNoguchi,\ufeffOhtaki\ufeff&\ufeffKamada,\ufeff2016).\ufeffIt\ufeffshows\ufeffthe\ufeffvertical,\ufefftriangular,\ufeff diamond\ufeffand\ufeffwedge\ufeffformation\ufefffrom\ufeffleft\ufeffto\ufeffright\ufeffin\ufeffFigure\ufeff8. To\ufeff confirm\ufeff the\ufeff practicability\ufeff of\ufeff this\ufeff approach,\ufeff this\ufeff study\ufeff validates\ufeff the\ufeff formation\ufeff from\ufeff the\ufeff horizontal\ufeffteam\ufeffinto\ufeffa\ufeffdiamond\ufeffand\ufefftrapezoidal,\ufeffand\ufeffeffect\ufeffof\ufefftransformation\ufeffis\ufeffindicated\ufeffin\ufeffFigure\ufeff9. The\ufeffexperiment\ufeffis\ufeffimplemented\ufeffwith\ufeffMatlab.\ufeffFirstly,\ufeff90\ufeffpositions\ufeffof\ufeffthe\ufefforiginal\ufeffformation\ufeffare\ufeff preset\ufeffand\ufeffassigned\ufeffto\ufeff10\ufeffgroups,\ufeffand\ufeffthe\ufeffcenter\ufeffof\ufeffevery\ufeffgroup\ufeffis\ufeffthe\ufeffleader.\ufeffThen,\ufeff90\ufeffpositions\ufeff of\ufeffgoal\ufeffformation\ufeffare\ufeffprovided\ufeffand\ufeffsplit\ufeffinto\ufeff10\ufeffzones\ufeffby\ufeffpolygons,\ufeffwhere\ufeffthe\ufeffnearest\ufeffpoint\ufeffto\ufeffthe\ufeff regional\ufeffcenter\ufeffis\ufeffthe\ufeffposition\ufeffof\ufeffleader.\ufeffBased\ufeffon\ufeffthe\ufeffgreedy\ufeffalgorithm,\ufeffthe\ufeffmapping\ufeffcoordinates\ufeffof\ufeff the\ufeffleader\ufeffand\ufeffmembers\ufefffrom\ufeffinitial\ufeffformation\ufeffto\ufeffthe\ufeffgoal\ufeffis\ufeffdrawn.\ufeffFinally,\ufeffcollision\ufeffavoidance\ufeffis\ufeff implemented\ufeffbased\ufeffon\ufeffthe\ufeffcontext\ufeffavoidance\ufeffrules\ufeff(as\ufeffshown\ufeffin\ufeffconversion\ufeff2). Figure\ufeff9\ufeffshows\ufeffthat\ufeffformations\ufeffstart\ufeffto\ufeffaggregate\ufeffat\ufeffconversion\ufeff1\ufeffwhere\ufeffthe\ufeffleaders\ufeffemerge;\ufeffand\ufeff the\ufeffgoal\ufeffformation\ufeffappears\ufeffin\ufeffconversion\ufeff2.\ufeffDuring\ufeffthe\ufeffwhole\ufefftransformation\ufeffprocess,\ufeffthere\ufeffis\ufeffno\ufeff collision\ufeffbecause\ufeffof\ufeffgaps\ufeffbetween\ufeffpoints.\ufeffIn\ufeffview\ufeffof\ufefftransformation,\ufeffthe\ufefffeasible\ufeffpath\ufeffis\ufeffachieved\ufeffby\ufeff greedy\ufeffalgorithm,\ufeffyet\ufeffnot\ufeffthe\ufeffoptimum." + ] + }, + { + "image_filename": "designv11_92_0000489_imece2009-10196-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000489_imece2009-10196-Figure7-1.png", + "caption": "Figure 7. Example of a student-produced microcar.", + "texts": [], + "surrounding_texts": [ + "Figure 1 illustrates one of the deliverables which is milestone #10 which details some of the tasks required for engineering analysis. In order to illustrate some of the quality learning by the students, exhibits of post-lab deliverables (milestones) are presented in Figures 2-7. These examples of student works are presented chronologically according to their production throughout the semester. Considering that these are students at an early stage of their college education, the relatively high quality of the technical content of these exhibits (Figures 2- 7) is quite evident. Figure 2 is an example of a studentproduced project schedule performed on MS Project. Figure 3 is an example of a student-produced free body diagram of car undergoing event 4: tug of war (see below). Figure 4 is an example of a student-produced systems engineering Pugh matrix (for evaluating potential drive system designs for performance in different contest events). Figure 5 is example of a student-produced measured igure 6 is an example of a student-produced CNC machined measurement of the drive motor\u2019s characteristic (torque vs. speed) curve. F microcar chassis. 3 Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/08/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use microcar chassis. . TEAM FUNCTIONS t of the course deliverables organized 4 The students produce mos into teams of four: manager, systems engineer, analyst, and CAD detailer. Team composition is based on functions with different skill sets as follows: Function I: The manager or the \u2018Big Wig\u2019 who forms the problem statement, does the scheduling and planning, manages manpower, conducts literature survey, and is responsible for overall report documentation. Function II: The systems engineer or the \u2018Big Picture Guy\u2019 bears primary responsibility for devising the contest winning strategy. This function defines the design parameters and translates requirements into engineering specification, prepare product configuration, identify potential scenarios, and do the hand sketches of the complete system assembly. Function III: The analyst or \u2018Mr. Brains\u2019 who is responsible for the engineering and mathematical solutions, sets the basic free body diagrams. Function IV: The detail designer or \u2018CAD guru\u2019 who is . TEAM FORMING nvolved in team forming. The first responsible for developing detailed designs using CAD tools and the detailed parts list documentation identifying the sources of components. 5 Two challenges are i challenge is individual students match to a suitable team function. Once individual assignments to the team functions are done, team composition becomes the challenge: do students pick their own teammates or should the four team members be assigned at random? For the task of assigning students to functions and during the first year ME Tools was introduced, team member selection was done based on student resumes submitted at the start of the semester and in which students were asked to identify their 1st and 2nd preferences. GA\u2019s assigned the students to the different functions based on these preferences and the rest of the information provided in the student\u2019s resume. More recently, a 20-question questionnaire was developed for identifying the student\u2019s functional preference based on their responses. Each function is targeted by a 5 focused questions dispersed throughout the questionnaire. Questions 1, 5, 9, 13, and 17; questions 2, 6, 10, 14, and 18; questions 3, 7, 11, 15, and 19; questions 4, 8, 12, 16, and 20 correspond to the functions of manager, systems engineer, analyst, and detail designer, respectively. These questions are shown in Figures 8-11. The highlighted answers reveal the more suitable response for the function of interest. Only two responses \u2018A\u2019 and \u2018B\u2019 are allowed. 1. In meetings, I generally tend to 4 Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/08/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use or each student, the responses are tallied. The responses to the he focused nature of each set of 5 questions is designed only . ASSESSMENT COMPONENTS AS MAPPED TO mbination of team grades as well as are ired ixed idu to nd al on (5%; Team. Benefits nding team based on ent F questions are totaled based on the difference between the \u2018A\u2019 and \u2018B\u2019 responses. Totaling the number of \u2018correct\u2019 responses allow us to determine the relative strength of the function preference (on a scale of 1-5). The higher the score the stronger is the preference for the function. For example, a student with 4B and 1A responses to any set of 5 questions in Figures 8-11 has a net response of 3B. For example, a student with 5A response to the group of questions in Figure 11 indicates a very strong tendency for performing detail tasks. Analyzing the responses methodically result in segregating the students along the four desired functions according to their highest function scores. T to reveal the affinity of the individual for a certain function. Therefore, it is quite possible for an individual to have strong preference for more than one function since having an affinity for one function does not necessarily exclude having other affinities for other functions. However, the questionnaire revealed that many students who displayed a preference for \u2018softer\u2019 functions such as the manager (and to a lesser extent systems engineer) indicate equally polarized response disfavoring \u2018hard\u2019 functions (analysts and detailed designer) and vise versa. Another revelation is that many students intuitively expressed a preference for a certain function prior to conducting the questionnaire which was found to correlate quite well to their questionnaire-revealed preference. Regarding team composition, once the function profile of each individual is determined, the students are randomly assigned to their respective teams. The intent is to mimic industrial occupational settings where employees do not usually get a chance to choose their team mates. 6 LEARNING METHODS The grading is based on co individual grades. Table 1 is a summary matrix showing the assessment components and their mapping to active, cooperative, collaborative, and problem-based learning methods. Specifically, these assessment components are: 1. Lecture Assignments (5%; Team): Students requ to do 5 homework sets related to lecture material. 2. Lab/shop Milestone Assignments (15%; M indiv al / team deliverables): Students are required to do twelve homework assignments related to lab/shop material. 3. Attendance (5%; Individual): Students are required atte l lectures, lab sessions, and team weekly meetings. 4. Project Notebook (5%; Individual but based student performance as a team member). 5. End-of-Term Team Presentation of cooperative learning for developing the interpersonal skills required for effective teamwork). Students are required to present at the end of the term to summarize their efforts. 6. End-of-Term Contest (25%; Team Grade depe on performance): Starting from common requirement statement, students (teams of four) start by identifying needs (functional requirements) and then devising practical solutions (Design Parameters) to fill those needs through (1) designing, (2) fabricating, (3) integrating, (4) and testing of an engineered product. Working from a common kit, the student goes about developing a functional engineering prototype. 7. Final Design Report (20%; Individual stud function in team): a major document delivered at the end of the term. The report assessment criteria for each team function are laid out according to strict assessment rules with the assessment components map precisely to the function\u2019s primary responsibilities for each function. Each category is given a corresponding symbol for identification and grading, e.g., M1-M5 for managers, and is given % of the report grade. 8. Final Exam (20%; Individual and contains elements learned during active learning at lab/shop) 5 Copyright \u00a9 2009 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/08/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use 6 Copyright \u00a9 2009 by ASME rding to Table 1 has Extensive evidence of student learning acco 8. CONCLUDING REMARKS been collected but was not included due to space limitations. The course introduces students to real world engineering skills and tools, so the students can evolve into professionals. The format and organization of this design course is such that it integrates the best of the four learning methods: active-, collaborative-, cooperative-, and problem-based learning. Students work in teams of four in order to design and construct an electric microcar which is to compete in an end-of-semester contest. Within the teams, four distinct functions are allocated: manager, systems engineer, analyst, and detail designer. Teamwork experience and communication skills are highly stressed and practiced." + ] + }, + { + "image_filename": "designv11_92_0002389_978-3-319-99010-1_8-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002389_978-3-319-99010-1_8-Figure5-1.png", + "caption": "Fig. 5. The membership function of De (k)", + "texts": [], + "surrounding_texts": [ + "The error e(k) and error change De(k) of stator current is the input of the observer, and then De\u00f0k\u00de \u00bc e\u00f0k\u00de e\u00f0k 1\u00de \u00f05\u00de e\u00f0k\u00de \u00bc I s \u00f0k\u00de Is\u00f0k\u00de \u00f06\u00de Where, k is switch state of inverter. I s \u00f0k\u00de is given stator current value, Is\u00f0k\u00de is actually measured stator current value [1] (Fig. 3). The language variable of e(k) and De(k) are defined by five fuzzy language variables {NB, NS, ZO, PS, PB}. The observer output is the stator resistance error DRs (k) [7]. DRs k\u00f0 \u00de \u00bc Rs k\u00f0 \u00de Rs k 1\u00f0 \u00de \u00f07\u00de Where, Rs (k) is the value of stator resistance this time; Rs (k \u22121) is the value of stator resistance last time [1]. The language variable DRs (k) is defined by five fuzzy language variables {NB, NS, ZO, PS, PB} [2]. 4.1 Fuzzification The regulator has two inputs, e(k), \u0394e(k) and for the fuzzyfied output R(k) as to Figs. 4, 5 and 6 [2]. The set of rules is described according to Mac Vicar with the format If- Thus, under the fuzzy rules table with two input variables according to [2] (Table 1). The choice of the inference method depends upon the static and dynamic behavior of the system to regulate, the control unit and especially on the advantages of adjustment taken into account. We have adopted the inference method Max\u2013Min because it has the advantage of being easy to implement in one hand and gives better results on the other hand [8, 12]. 4.3 Defuzzyfication The most used defuzzification methods is that of the center of attraction of balanced heights, our choice is based on the latter owing to the fact that it is easy to implement and does not require much calculation [12]." + ] + }, + { + "image_filename": "designv11_92_0002894_ihmsc.2018.10168-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002894_ihmsc.2018.10168-Figure4-1.png", + "caption": "Figure 4. simplified model of rotating shaft", + "texts": [ + " The dynamic model of the helicopter in the pitching axis is as follows: ( )bffhp VVKpLJ p \u2212= \u2022\u2022 cos 4 Among them, J p the moment of inertia for the pitch axis; The deflection Angle of pitch axis p ; K f Is the motor constant; The length of the pitch axis to the propeller Lh . C. Rotating Shaft Modeling When the thrust of the helicopter is not equal to the thrust produced by the front and rear propellers, it will naturally produce a lateral component to push the lateral movement of the fuselage, which is the power source of the rotation of the rotating shaft. Fig. 4 is a simplified model of the rotating shaft. The dynamic equation of the rotating shaft reflects its coupling relation with the pitching axis: gat FpLJ t sinsin \u03b5\u2212= \u2022\u2022 5 Where, J t the moment of inertia for the rotation axis; t For rotation axis deflection Angle; F g Effective gravity for the helicopter. Considering the actual situation of helicopter flight, the value of pitch angle of P will reduce the ride comfort, and even affect the flight safety, so the p value should be controlled in a small range, and can be type (5) for linear approximation: gat FpLJ t \u03b5sin\u2212= \u2022\u2022 6 Select the helicopter height Angle \u03b5 , pitch Angle p , rotation Angle t , their differentia \u03b5 \u2022 p \u2022 t \u2022 , and the integral of the height Angle and the rotation Angle , and establish the eight dimensional state space equation: \u23a5 \u23a6 \u23a4 \u23a2 \u23a3 \u23a1 + \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 = \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u222b \u222b \u2022 \u2022 \u2022 \u2022\u2022 \u2022\u2022 \u2022\u2022 \u2022 \u2022 \u2022 V VB t t p t p A t t p t p b f \u03b5 \u03b5 \u03b5 \u03b5 \u03b5 \u03b5 7 Among them \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 = 00000100 00000001 000000 sin 0 00000000 00000000 00100000 00010000 00001000 J FL A t ga\u03b5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a5 \u23a6 \u23a4 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a2 \u23a3 \u23a1 \u2212 = 00 00 00 coscos 00 00 00 J KL J KL J LLp J LLp B p fh p fh fafa \u03b5\u03b5 According to the actual situation of the helicopter flight model in the process of using model switching, but not in the traditional sense of the weighted model theory, according to the specific location of the helicopter, the model parameters for the system, and to design multiple LQR controller, each controller is respectively set in line with the current situation of the Q and R matrix values, control volume the most direct and effective, the helicopter tracking is fast and accurate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002171_978-3-319-49574-3_12-Figure12.28-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002171_978-3-319-49574-3_12-Figure12.28-1.png", + "caption": "Fig. 12.28 Example showing test point positions (TP1, TP2)", + "texts": [ + " Where high values are found, the joint should be remade. When complete, move on to record measured values. If the measured value is less than the calculated value (acceptable), the value should be recorded. If the measured value is greater than the calculated value, measure each busbar joint to see which joint is reading higher than 10 \u03bc\u03a9. Remake the busbar joint and repeat the resistance measurement. In the example above, seven joints are between the two test points, each joint should be 10 \u03bc\u03a9. The seven joints shown in Fig. 12.28 would be calculated as: 7 x 10 \u03bc\u03a9: \u00bc 70 \u03bc\u03a9, therefore, measurement between TP1 and TP2 70 \u03bc\u03a9: Solid-core post insulators are used in the assembly of disconnectors described in Sect. 12.3 above and for supportingHVelectrical conductors in substations and incorporated in medium-voltage equipment. They are a simple product but extremely effective at performing such a function. Hollow insulators are used for supporting electrical equipment and to provide an internal chamber for connections, either electrical for items such as bushings, current transformers, voltage transformers, and power transformers or mechanical, typically for the drive rod of circuit breakers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003577_9780470459300.ch16-Figure16.16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003577_9780470459300.ch16-Figure16.16-1.png", + "caption": "FIGURE 16.16 Absorption mechanisms inside the keyhole. (a) Inverse bresmsstrahlung absorption. (b) Fresnel absorption. (From Beyer, E., Behler, K., and Herziger, G., 1988, SPIE, High Power Lasers and Laser Machining Technology, Vol. 1020, pp. 84\u201395.)", + "texts": [ + " The lower part is nearly parallel sided, with a gradual narrowing from top to bottom. On top of that is a region of roughly semicircular shape, whose width is greater than that of the parallel-sided section (Fig. 16.15). 16.3.2.1 Power Absorption in the Keyhole Absorption of beam energy in the keyhole occurs by two principal mechanisms: 1. Inverse bremsstrahlung of the plasma. 2. Fresnel absorption (direct deposition) at the cavity walls. Both forms of absorption normally occur simultaneously within the keyhole, and are illustrated in Fig. 16.16. Their relative magnitudes are still the subject of research. The absorbed beam becomes entrapped in the keyhole. This results in significantly higher absorption in deep penetration welding compared to conduction mode welding. Inverse Bremsstrahlung Absorption With inverse bremsstrahlung absorption, a portion of the vaporized metal or shielding gas is ionized by the high incident energy to form a plasma that may absorb part of the incident beam by the inverse bremsstrahlung effect inside the keyhole. Bremsstrahlung is the electromagnetic radiation produced by the sudden retardation of a charged particle in an intense electric field. Thus in simple terms, inverse bremsstrahlung is the absorption of electromagnetic radiation as a result of the sudden acceleration of a charged particle, or absorption of a photon by an electron in the field of a nucleus. The absorption thus depends on the presence of free electrons and occurs within the plasma, which consists of ionized particles (Fig. 16.16a). The absorption depends on temperature, partly because the electron density varies with temperature. As the vapor is heated in the presence of an intense beam, the absorption coefficient increases. Since the center of the keyhole is much hotter than the walls (which are at the vaporization temperature), most of the beam energy is absorbed within the cavity, and then coupled to the keyhole wall through conduction and radiation. For inverse bremsstrahlung absorption, the corresponding absorption coefficient, \u03b1p, is given by equation (14.13). Fresnel Absorption The mechanism of Fresnel absorption results in the beam being spread along the entire keyhole length. It results from multiple reflections at the keyhole walls (Fig. 16.16b), where the beam energy that enters the vapor cavity is reflected repetitively at the walls of the keyhole until all the energy is eventually absorbed. For most applications, specular reflection (mirror-like) is assumed to be predominant, with diffuse reflection (similar to reflection from a rough surface) being neglected. Such multiple reflections, along with scattering processes, tend to randomize phase relations that may have existed in the original beam. The absorption coefficient for Fresnel absorption can be estimated by first determining the coefficient for Fresnel reflection, RF, which for a circularly polarized light, is given by RF(\u03b8) = Rp(\u03b8) + Rs(\u03b8) 2 (16" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002890_j.matpr.2018.08.148-Figure8-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002890_j.matpr.2018.08.148-Figure8-1.png", + "caption": "Fig. 8. Displacement contour lines for the maximal deflection of the rubber bumpers.", + "texts": [ + " Boundary conditions, enabling this type of the analysis of the structure were used. The loads acting between the leaves of the spring was considered by applying a proper contact connection. In order to verify a virtual model, comparison of the numerical and experimental tests was made. The authors compared obtained strains, displacements and the reaction forces. There was a broad convergence of the results (fig. 7). After adjusting the numerical model, the influence of the mounted rubber bumpers on the stress values was examined. Figure 8 presents the analysed object in the initial position and in the moment of maximal deflection of the rubber rolls. Obtained from the numerical simulation contour lines of the stress confirmed, that the cracks appeared on the object in the spots where the maximal stresses occurred. 26764 Sta\u0144co M., Iluk A., Dzia\u0142ak P./ Materials Today: Proceedings 5 (2018) 26760\u201326765 The crack of the shortest leaf occurred in the area near the mounting of a spring. Figure 9 presents distribution of the maximal principal stresses \u03c31 in the fractured leaf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000274_amr.44-46.127-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000274_amr.44-46.127-Figure2-1.png", + "caption": "Fig. 2 Lumped mass model of shaft element m", + "texts": [ + " The rigid disk 3 is the equivalent disk of the high-pressure turbine. The high pressure rotor and the radial driveshaft (shaft 2) are geared by a spiral bevel gear pair. When the engine is started, shaft 2 acts as the driving shaft; when the engine is in cruises, shaft 1 acts as the driving shaft. Dynamic Model of the Coupled Rotor System There are two shafts in the coupled rotor system. Each shaft is regarded as a shaft element.The two shaft elements are coupled by the coupled matrix of the spiral bevel gear pair. As shown in Fig.2[9], a local reference frame m m m mo x y z is used to describe the motions of the shaft m ( 1,2m = ). The mz axis is collinear with the axis the shaft m, the ,m mx y plane perpendicular to the mz axis. The shaft 1 is divided into 17 segments and the shaft 2 is divided into 11 segments (Fig.1). Each segment is composed of one massless elastic short shaft and one rigid disk that is the equivalent disk of the inertial parameters on the node between the two segments. As shown in Fig.2 , jl is the length of segment j of the shaft element m, jd is the diameter of segment j of the shaft element m, jEI is the bend stiffness of segment j, jK\u03b8 is the torsional stiffness of segment j, jm is the mass of the equivalent disk j , ( ), ,ijJ i x y z= is the moment of inertia, ( ), ,jpqd p q x y= is the damping of the bearing, and ( ), ,jpqk p q x y= is the stiffness of the bearing. In the local reference frame m m m mo x y z , the free vibration equations set of the node j can be obtained as 1 1 1 1 { } 0j j jj j j j j j a j j c jb l d + + + \u2212+ \u2212 \u2212 + + + =&& &M X C X K X K X K K K X (1) where yj x z j m m J J J = M 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 xx xy yx yy j z z j d d d d J J \u03c9 \u03c9 = \u2212 C 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 xx xy yx yyj l j k k k k = K 1 1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 j a j A B A B B C B C K\u03b8 + + \u2212 \u2212 \u2212= \u2212 K 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 j b j A B A B B C B C K\u03b8 \u2212 \u2212= \u2212 \u2212 K 1 1 0 0 0 0 0 0 0 2 0 0 0 0 2 0 0 0 0 0 j c j A B A B B C B C K\u03b8 + + = K 0 0 0 0 0 0 0 2 0 0 0 0 2 0 0 0 0 0 j d j A B A B B C B C K\u03b8 \u2212 \u2212 \u2212= \u2212 K { , , , , }T j j j j j j mx y \u03d5 \u03c8 \u03b8=X is the node j displacement of the shaft element m in the local reference frame m m m mo x y z " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003738_978-3-030-04975-1_7-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003738_978-3-030-04975-1_7-Figure1-1.png", + "caption": "Fig. 1. Cycloidal gear transmission", + "texts": [ + " In planetary gear transmissions, cycloidal curves are successfully used as a gear profile [10]. A distinction is made between two versions of those transmissions called cycloidal gears [9]. In both types of the transmissions, the cycloidal gear collaborates with the rollers making the teeth of the other gear. In one of them, the epicycloidal curve makes the contour of the external gear, whereas in the other one, the hypocycloidal curve is the outline of the internal gear. The transmission with the hypocycloidal gearing is shown in Fig. 1 [8]. The cycloidal gears and the rollers form the internal gearing. The cycloidal transmissions are mainly used in drive systems as reducers. A cycloidal gear transmission is a rolling transmission in which all geometrically connected elements are moved by rolling motion (Fig. 2). This results in a maximum reduction of losses caused by friction. Three rolling pairs can be distinguished: \u2022 planetary gear (1) with rollers (2) collaborates with fixed body (6) with the hypocycloidal profile; rollers (2) rotate around their own axis at speed xr; \u2022 planetary gear (1) collaborate with the central bearing (of the eccentric) (3); the gear is fixed eccentrically on the input shaft, moving at angular speed xh; \u2022 planetary gear (1) collaborates with sleeve (5), together with pin (4) moving at speed xt in the hole of the planetary gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002529_012004-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002529_012004-Figure10-1.png", + "caption": "Figure 10. for a selected binder and row of warp tows", + "texts": [ + " Again, the normalized stresses shown are in the local coordinate system. Overall, Figure 9 shows that the warps and binders experience severe transverse tension ( ), while the wefts experience severe longitudinal shear ( and ). The most severe stress in the textile is in the binders and warps. For both types of tows, the local y-axis is closely aligned with the global y-axis, which is the direction of the load for this configuration. To illustrate where the peak stresses occur in the textile, Figure 10 shows for one binder and row of warp tows, along with the local coordinate system for a point in the binder. The figure shows that the stress concentrations form where warps and binders come closest, such as point A in Figure 10. Additionally, stress concentrations form where the binder begins or ends traveling through the thickness of the textile, such as points B in Figure 10. 13 1234567890\u2018\u2019\u201c\u201d IOP Conf. Series: Materials Science and Engineering 406 (2 18) 012004 doi:10.1088/1757-899X/406/1/012004 14 1234567890\u2018\u2019\u201c\u201d IOP Conf. Series: Materials Science and Engineering 406 (2 18) 012004 doi:10.1088/1757-899X/406/1/012004 In summary, the transverse normal stress was the most severe component for both configurations. In both cases, the peak occurred within the tows that are perpendicular to the load. However, when the load is along the global x-axis, the stress concentrations only form within the wefts, but when the load is along the y-axis, the stress concentrations form within the binders and warps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003679_978-3-030-11827-3_9-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003679_978-3-030-11827-3_9-Figure2-1.png", + "caption": "Fig. 2 General sight of the experimental installation", + "texts": [ + " In this study, acquisition of vibratory signals was performed during machining using a Br\u00fcel & Kjaer type 4524B triaxial piezoelectric accelerometer placed in the area closest to the cutting sector to record the intensity of the accelerations in real time in the three main directions (x, y, z). The frequency band selected varies between (0\u201312, 800 Hz) where each signal contains 16,384 points. The measurement results were stored directly on the PC using the analyzer acquisition system, driven by Br\u00fcel & Kjaer\u2019s Pulse Lab shop\u00ae software (Fig. 2). The development of the flank wear on the cutting insert is measured after each test; the cutting insert is removed from the tool holder, cleaned and then placed on the table of the optical microscope Standard Gage type Visual 250. Having an optical enlargement from 0.7\u00d7 to 4.5\u00d7 the actual size. The Visual Gage software allows any operator as well as the metrologist to measure quickly and accurately most geometric elements. Three experimental companions of 25 trials have taken several values of the flank wear and tear of the three plates studied, from the new state to the end of the lifetime (see Table 1)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001766_978-981-10-7212-3_3-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001766_978-981-10-7212-3_3-Figure4-1.png", + "caption": "Fig. 4 The position mounting diagram of sensors", + "texts": [ + " The amount of absorption is related to the concentration of the gas and is measured by a set of optical detectors and suitable electronic systems. The change in the intensity of the absorbed light is measured relative to the intensity of light at a non-absorbed wavelength. The microprocessor computes and reports the gas concentration from the absorption (Fig. 3). The gas detection system is equipped with eight infrared absorption CO2 sensors which are mounted in the main controller of micro-drone. The structure of this system is described in Fig. 4. In this plane, these eight gas sensors are connected by two carbon fiber rods which are, respectively, arranged in mutually orthogonal. In this paper, the 3D gas tracking algorithm is proposed to follow the plume by quality of dates from this gas detection system. According to different concentrations of gas sensors perception in different positions, the direction of following the plume can be got. The body coordinate system \u00f0x, y, z\u00de is transformed into the ground coordinate system \u00f0X,Y ,Z\u00de when the aerial vehicle in steady flight" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002805_detc2018-85407-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002805_detc2018-85407-Figure4-1.png", + "caption": "Figure 4. Dynamical model of the steer system", + "texts": [], + "surrounding_texts": [ + "method is simple and there are few parameters influencing the accuracy of estimation. In this paper, an innovative method to estimate the road adhesion coefficient which using the dynamic characteristics of the tire rebound after steering when the vehicle is stationary is proposed. A vehicle equipped with steerby-wire system is tested in this paper. After the steering wheel has turned to a certain angle, releasing the steering wheel suddenly will result in front wheel rebound to a certain angle. Figure 1 shows the diagram of the steer-by-wire system used in the experiment. Figure 2 shows the experiment vehicle. For this vehicle, the front wheel turning angle is proportional to the steering wheel turning angle. Figure 3 shows the steering wheel turning angle obtained by the steering wheel angle sensor in this experiment. And the experiment road surface is snow surface in figure 3. (The experiment is conducted in a professional experiment site, and the road adhesion coefficient is uniform for the same kind of road surface.) The red curve in the figure represents the steering wheel turning angle, and the green curve in the figure represents the output torque of steering motor in SBW system. Obviously, when the steering torque suddenly changes to zero, the steering wheel will turn backward for a certain angle, which means the front wheel will rebound. When the steer wheel turns the same angle on different road whose adhesion coefficients are different, the front wheel rebound angles are different.\nFigure 1. The diagram of the steer-by-wire system used in the\nexperiment\nThe detailed dynamical model of the front wheel rebound process is established in this paper. And the rebound is divided into positive rebound and reverse rebound, which is showed in figure 3. The direction of positive rebound is as same as the direction of self-aligning moment. And the direction of reverse rebound is opposite to the direction of self-aligning moment. The initial process, the middle process and the final process of wheel is modeled. The Luenberger reduced-order disturbance observer is established based on the middle process of the positive rebound. The friction moment between the road and wheel is estimated using this observer. Final, the road adhesion coefficient which equals to road friction coefficient is obtained. The estimation results are quite reasonable, which demonstrate the validity of this method.\n2 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "Before releasing the steer torque\nFigure (4) shows dynamical model of the steer system. Equation (1) and (2) describe the model. After the steering wheel has turned to a certain angle, the resultant force is zero in steer system. At this time, the elastic deformation of front wheel occurs. And the center of gravity of the vehicle is raised owing to kingpin angle.\n0 = \ud835\udc47\ud835\udc60\ud835\udc61\ud835\udc52\ud835\udc52\ud835\udc5f \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 \u2212 \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 (1)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 = \ud835\udc58\ud835\udc61 \u2219 (\ud835\udeff\ud835\udc610 \u2212 \ud835\udeff\ud835\udc500) (2)\nIn Eq. (1), \ud835\udc47\ud835\udc60\ud835\udc61\ud835\udc52\ud835\udc52\ud835\udc5f represents the steer torque which is produced by steer motor acting on the steering column, \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 represents the torque produced by the viscous friction in the steer system, \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 represents friction torque produced by the static friction force between road and tire. \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b represents the self-aligning torque acting on the kingpin.\nIn Eq. (2), \ud835\udc58\ud835\udc61 represents elastic coefficient of the front tire, \ud835\udeff\ud835\udc610 represents the front wheel turning angle at this time (\ud835\udeff\ud835\udc610 can be regarded as the turning angle of the rim and the turning angle of upper part of tire), \ud835\udeff\ud835\udc500 represents the turning angle of the tire part contacting with the ground (\u201ccontact part\u201d will be used to represent this part in the following).\nAfter releasing the steer torque\nThe initial process of the wheel reverse rebound\nDuring this process, the tire part contacting with ground doesn\u2019t turn. Meanwhile, the upper part of tire which connect to the rim rebounds, and the steer wheel turns backward. Equation (3-5) describe the dynamical model. The steering system can be equivalent to the following diagram showed in figure (5).\nreverse rebound\n\ud835\udc3d\ud835\udc60?\u0308?\ud835\udc60 + \ud835\udc3d\ud835\udc611?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc61?\u0307?\ud835\udc61 + \ud835\udc3e\ud835\udc61\ud835\udeff\ud835\udc61 + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \u2212\ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 (3)\n\ud835\udeff\ud835\udc60 = \ud835\udc58\ud835\udc60 \u2219 \ud835\udeff\ud835\udc61 (4)\n\ud835\udc3d1 = \ud835\udc58\ud835\udc60 \u2219 \ud835\udc3d\ud835\udc60 + \ud835\udc3d\ud835\udc611 (5)\nIn above equations, \ud835\udc3d\ud835\udc60 represents the equivalent moment of inertia of the system except front wheel, \ud835\udc3d\ud835\udc611 represents the moment of inertia of the upper part of tire, \ud835\udc35\ud835\udc61 represents the damping coefficient of the tire, \ud835\udeff\ud835\udc60 represents the steering wheel turning angle, \ud835\udeff\ud835\udc61 represents the front wheel turning angle, \ud835\udc58\ud835\udc60 is a constant, \ud835\udc3d1 represents the equivalent moment of inertia of the steer system, \ud835\udc35\ud835\udc60 represents the damping coefficient of the supporting mechanism in steer system. The initial value of \ud835\udeff\ud835\udc61 is \ud835\udeff\ud835\udc610.\nThe final process of the wheel reverse rebound\nEquation (6) and (7) describe the dynamical model. And the\nfigure (6) shows the diagram of the model.\nreverse rebound\n0 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 (6)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 = \ud835\udc58\ud835\udc61 \u2219 (\ud835\udeff\ud835\udc611_\ud835\udc5f \u2212 \ud835\udeff\ud835\udc500) (7)\n3 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use", + "In above equations, \ud835\udeff\ud835\udc611_\ud835\udc5f represents the steering front\nturning angle at this time.\nThe initial process of the wheel positive rebound\nThe upper part of tire rebound because of the elastic force. But the elastic force is less than the maximum static friction. So the tire part contacting with the ground doesn\u2019t turn. Equation (8) and (9) describe the dynamical model. And the figure (7) shows the diagram of the model.\npositive rebound\n\ud835\udc3d1?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc61?\u0307?\ud835\udc61 + \ud835\udc3e\ud835\udc61\ud835\udeff\ud835\udc61 + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 (8)\n(\ud835\udc35\ud835\udc61?\u0307?\ud835\udc61 + \ud835\udc3e\ud835\udc61\ud835\udeff\ud835\udc61) \u2264 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 (9)\nEquation (10) and (11) describe the dynamical model of the\nfinal time of this process.\n0 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60 \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 (10)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 = \ud835\udc58\ud835\udc61 \u2219 (\ud835\udeff\ud835\udc611_\ud835\udc53 \u2212 \ud835\udeff\ud835\udc500) (11)\nIn above equations, \ud835\udeff\ud835\udc500 represents the turning angle of the tire part contacting with the ground at this time, \ud835\udeff\ud835\udc611_\ud835\udc53 represents the front wheel turning angle at this time.\nThe middle process of the wheel positive rebound\nEquation (12) describes the dynamical model. And the figure\n(8) shows the diagram of the model.\npositive rebound\n{ \ud835\udc3d1 \u2219 ?\u0308?\ud835\udc61 + \ud835\udc35\ud835\udc61 \u2219 (?\u0307?\ud835\udc61 \u2212 ?\u0307?\ud835\udc50) + \ud835\udc3e\ud835\udc61(\ud835\udeff\ud835\udc61 \u2212 \ud835\udeff\ud835\udc50) + \ud835\udc35\ud835\udc60 \u2219 ?\u0307?\ud835\udc61 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc59\ud835\udc5c\ud835\udc60\ud835\udc60\n\ud835\udc3d\ud835\udc612 \u2219 ?\u0308?\ud835\udc50 + \ud835\udc35\ud835\udc61 \u2219 (?\u0307?\ud835\udc50 \u2212 ?\u0307?\ud835\udc61) + \ud835\udc3e\ud835\udc61(\ud835\udeff\ud835\udc50 \u2212 \ud835\udeff\ud835\udc61) = \u2212\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50\n(12)\nIn above equation, \ud835\udc3d\ud835\udc612 represents the moment of inertia of the tire part contacting with the ground, \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50 represents the friction torque produced by the sliding friction force between road and tire.\nThe final process of the wheel positive rebound\nEquation (13) and (14) describe the dynamical model of the final time of this process. And the figure (9) shows the diagram of the model.\npositive rebound\n0 = \ud835\udc47\ud835\udc4e\ud835\udc59\ud835\udc56\ud835\udc54\ud835\udc5b \u2212 \ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 (13)\n\ud835\udc47\ud835\udc53\ud835\udc5f\ud835\udc56\ud835\udc50_\ud835\udc60 = \ud835\udc58\ud835\udc61 \u2219 (\ud835\udeff\ud835\udc612_\ud835\udc53 \u2212 \ud835\udeff\ud835\udc501) (14)\nIn above equations, \ud835\udeff\ud835\udc501 represents the turning angle of the tire part contacting with the ground at this time, \ud835\udeff\ud835\udc612_\ud835\udc53 represents the front wheel turning angle at this time.\n4 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/07/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use" + ] + }, + { + "image_filename": "designv11_92_0000302_pi-b-2.1962.0194-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000302_pi-b-2.1962.0194-Figure1-1.png", + "caption": "Fig. 1.\u2014Construction showing the Lorentz transformation of time and distance as experienced by two observers of a mutual velocity c/2.", + "texts": [ + " (2), x \u2014 x'{\\ \u2014 v' These are Lorentz transformations, and the apparent contradiction which they express is sometimes regarded as a paradox; but the mental difficulty in accepting their logical consequence stems simply from an unwarranted\u2014though intuitive\u2014 denial that the velocity of light is independent of all other movement. Once this is accepted as an experimental fact, both length and time lose any claim to absolute significance and depend upon whether the system to which they apply is seen in motion or at rest. Length and time become relative. To grasp how these relative qualities are endowed by motion, it is helpful to construct a time/distance diagram which can show both S and S' views of a linear movement, as Professor Peierls has done.2 In Fig. 1 the system S is shown with orthogonal axes, and the S unit of distance xl is chosen to be c times the S unit of time t^ [244] RUGBY SUB-CENTRE: CHAIRMAN'S ADDRESS: MAGNETIC FORCE AND THE MOVING CHARGE 245 This means that the velocity c would lie at 45\u00b0 between the orthogonal axes. A point O', taken as the origin of a moving system S', is considered to move along x with a velocity v assumed, for illustration, to be as high as c/2. Its path in S becomes its S' ordinate, x' = 0, and is the line OA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001076_icmel.2008.4559321-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001076_icmel.2008.4559321-Figure3-1.png", + "caption": "Fig. 3. The elements of a contact", + "texts": [ + " It is a system of rigid bodies connected by one DoF joints which has no other constraints given in advance. No other force is acting upon it except gravity and the inter-segment forces coming from its own actuators, hence the name \u2013 flier. The flier alone, without any interaction with its environment wouldn\u2019t be enough for modeling a diversity of situations. Therefore, contacts with other objects are also possible. With flier mechanisms interacting through contacts, almost every sit- uation can be easily modeled. The basic elements of a contact are shown in Figure 3. Generally, a contact joins two bodies togerher and as a result both of them lose a few DoFs. In order to describe a contact, a contact point has to be introduced (Figure 3) and a so-called functional coordinate system placed in it. Its orientation should be the most apropriate for the analyzed contact situation. According to this coordinate system, 6 DoFs can be defined: 3 translations along each axis and 3 rotations around each axis. In a particular contact, the bodies in contact are unable to change their relative position described by these 6 DoFs \u2013 some DoFs are lost. But there are situations when not all 6 DoFs get blocked, e.x. writing on a sheet of paper \u2013 the pencil is in contact with the paper, but still can (and should) move along its surface, therefore only 4 DoFs are blocked" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001764_978-3-319-89480-5_7-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001764_978-3-319-89480-5_7-Figure5-1.png", + "caption": "Fig. 5 Methodology of process model", + "texts": [ + " An error in line spacing is also correlated with increased roughness of the top surface upon melting, which increases the propensity for build defects [2, 3]. The process model developed here predicts the mean track dimensions as a function of power, speed (and other process parameters), which makes it feasible to compute the line spacing based on a geometric argument that equates the overlap area between the adjacent tracks and the dip area such that when the two tracks fuse, the resulting top surface is close to being a flat even surface [3]. The modeling methodology is shown in Fig. 5. The process model was modified to consider a common defect in AM where the area fill and the contour tracks are not properly merged, which results in lack-of-fusion and/or subsurface porosity defects as shown in Fig. 6. The model was developed to select the correct contour power and velocity for Alloy 230 material such that defect regions at the ends of hatch tracks were re-melted and filled by the melt during the contour passes. The temperature map and melt pool dimensions on the contour pass predicted by the model are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002845_ipec.2018.8507656-Figure13-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002845_ipec.2018.8507656-Figure13-1.png", + "caption": "Fig. 13.Experimental setup", + "texts": [ + " SIMULATION RESULTS To prove the proposed current profiling being effective for torque ripple minimization, some simulation based on 2D-FEA is executed. Fig. 12 demonstrates torque and current waveforms under the conventional control and the proposed control methods for 100%, 200% and 300% load conditions at 100r/min. As can be seen in the figure, the torque ripple is minimized. VI. EXPERIMENT RESULTS To get a magnetizing characteristic of tested motor and evaluate a performance of torque ripple minimization, the experimental equipment is set up as can be seen Fig. 13. Fig.13 (a) is a equipment to obtain magnetizing curve. To obtain magnetic flux linkage data at fixed position, the voltage and current data through the pulse voltage injection is needed. To fix the position exactly for measurement, rotary encoder (Tamagawa Seiki, TS5667N420, 131072 ppr, 0.0027/pulse) and stepping motor (oriental motors, UPH599HG2-B2, 0.0072deg/pulse) with harmonic gear (oriental motors, parking torque 360kg.fcm, gear ratio 1:100) to have high resolution are used. Fig. 13 (b) is a equipment to take torque ripple experiment. The PE-Expert III (Myway Plus Corporation) was used as a controller to generate gate signals for the asymmetric Hbridge converter with a PWM frequency of 10 kHz. For position sensing, a encoder (same with above) was attached. As a torque meter and load, torque detector (ONO SOKKI, MT-6254A, Max. torque : 5Nm, resolution : 0.01Nm) consisting of a geared DC motor and torque sensor was coupled to the tested SRM and used as a constant speed dynamo" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002300_med.2018.8442905-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002300_med.2018.8442905-Figure2-1.png", + "caption": "Figure 2. The geometric structure of the quadrotor", + "texts": [ + " There is a gain-scheduled controller with structure (3), if there are checked matrices; R, S, L3 and J3 solved (5)-(8). (5) (6) (7) (8) The theorem was proofed in [22] for LTI systems and adopted for LPV systems in [18]. We assume a given solution of R, S, L3 and J3 of (5)-(8); in last we can get the controller\u2019s matrices with structure (9) and the solution of the LMI is shown in (10). (9) (10) Where: (11) And: (12) III. QUASI-LPV MODELING OF QUADROTOR-ATTITUDE The mathematical model of the quadrotor was generated by both of techniques; Euler-Newton [23] and EulerLagrange [14]. Fig.2 presents a general view of the quadrotor, which has four rotors, and where each two rotors turn on the same direction. We can find two kind of configurations, the cross (\u00d7) configuration and the plus (+) configuration, that is chosen in this paper. In this paper, we focus on the attitude subsystem, where the model has three degrees of freedom; roll, pitch and yaw; Six states; which are three angles with their derivatives , and three inputs (13): (13) Where : r= 1- 2+ 3- 4 (14) And: (15) Table I" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001861_978-3-319-79111-1_42-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001861_978-3-319-79111-1_42-Figure1-1.png", + "caption": "Fig. 1 Architecture of the PM", + "texts": [ + " The MP has only two degrees of freedom (DOF), two rotations allowed by the U joint of the passive constraining leg, fact that allows the orientation of a solar panel in order to match the relative sun angles i.e. azimuth and elevation. The number of degrees of freedom (DOF) M has to be lower than the number of actuators N with a constant factor of 1 due to the fact that the actuators (the Festo Fluidic muscles) are able to exert force only on contraction. Therefore, the PM is similar to the cable robots. The architecture is presented in the Fig. 1a. The topology of the PM is presented in the Fig. 1b (with only one UPS open loop). In order to optimize the geometrical lengths of the linkages of the PM with regard to the WS, the kinematics has to be well defined. Since the purpose of the PM is to orientate a solar panel, the desired WS of the PM is given as a set of angles (\u03b31, \u03b32) of the U joint A (Fig. 1b). These angles are the input in the inverse kinematic; the matching lengths of the fluidic muscles are computed. The Fig. 1b illustrates the topology of the PM. The other notations from the Fig. 1b represent: \u2022 O0x0y0z0 and Pxyz are the coordinate systems attached to the FP and MP; \u2022 O0 and P have been chosen as the centers of the circles containing the U and S joints attached to the FP and MP due to the fact that the complexity of the mathematical model is reduced by the symmetry; \u2022 l1 and l2 are the lengths of the elements of the constraining leg; \u2022 \u03b1i and \u03b2i are the angles of positioning the U and S joints on the FP and MP. Each two consecutive U or S joints are positioned 2\u03c0/3 radians apart (due to symmetry reasons)", + " As presented in [8], the position of the sun relative to the Earth can be defined by two angles, the elevation \u03b1s and the azimuth \u03b3s, which have to define also the orientation of the solar panel from the MP. A complete set of angles \u03b3s and \u03b1s has been computed using the approach from [8]. for the city Cluj-Napoca, Romania between 6 a.m. and 17 p.m. for several days between WiS and SuS. This hour interval assures positive elevation (i.e. above the horizon, therefore visible sun) even in the winter solstice. This set of angles illustrated in Fig. 2 is the input into the optimization problem. In ideal case, the PM presented in the Fig. 1 is able to cover each combination of azimuth\u2014elevation angles (\u03b3s, \u03b1s). An optimization problem computes the optimal values of a design vector. For the case of the problem from this paper, the design vector x from the Eq. (3) is represented by the geometrical characteristics of the robot. Therefore, the design vector contains the lengths of the passive constraining leg (l1 and l2 measured in m), the radii of the FP and MP (R and r measured in m) and the angles of positioning of the first U and S joints (\u03b11 and \u03b21 measured in radians) from the FP and MP, the other angles being evaluated from the first ones, as presented in the Sect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001599_ibcast.2018.8312301-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001599_ibcast.2018.8312301-Figure1-1.png", + "caption": "Fig. 1: AUV Model shown in body and earth fixed coordinates [4]", + "texts": [ + " ' '\u03b7 is vector of earth fixed frame quantities and ' '\u03bd is vector of body fixed frame quantities as described below [ ]Tx y z\u03b7 \u03c6 \u03b8 \u03c8= (2) [ ]v Tv w p q r\u03bc= (3) The earth fixed frame vector ' '\u03b7 in (2) contains AUV\u2019s positions and their respective orientations. And the body fixed frame vector ' 'v in (3) contains AUV\u2019s linear and angular velocities in x, y and z directions. The control forces acting on AUV can be described as, [ ]( ) ( ) ( )s rf f f n\u03c4 \u03b4 \u03b4= (4) here \u2018n\u2019 represents the propeller revolutions of AUV, the steering angle is represented by ' 's\u03b4 and the rudder angle The quantities related with position and motion of AUV body and earth fixed coordinates are explained in Fig.1. For simplification purposes, the earth fixed frame vector ' '\u03b7 is transformed to body fixed frame of reference applying Jacobian transformation as defined below, ( )J v\u03b7\u03b7 = (5) A 6DOF model of AUV is defined below ( ) ( ) ( ) ( ) Mv C v v D v v g J v \u03b7 \u03c4 \u03b7 \u03b7 + + + =\u23a7 \u23ab \u23aa \u23aa \u23a8 \u23ac \u23aa \u23aa=\u23a9 \u23ad (6) Notions used are adapted from SNAME [6]. A generalized mathematical model of AUV is presented by Thor I. Fossen [1]. ( ) ( )X f x g x u= + (7) Since stabilization problem of steering plane control of AUV is under consideration, its motion is decoupled in steering plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003313_imece2018-86323-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003313_imece2018-86323-Figure3-1.png", + "caption": "Fig 3: Slip and Steering Angle Diagram", + "texts": [ + "org/about-asme/terms-of-use Tan avg L R (2) By using Newton\u2019s Second Law steady-state relation for three axle vehicle was derived from cornering equations [13]: Rear axle angle relative to datum rear = ( ) r b c R (3) First axle angle relative to datum = 1 ( ) f a R (4) Second axle angle relative to datum = 1 ( ) f a R (5) Steering angle first axle: 1 1 ( ) f r a b c R (6) Steering angle second axle: 2 2 ( ) f r c R (7) 2 1 1rS (8) 1 1 2 2 0f yf f yf r yra b b cC C C (9) 2 1 1 2 2f yf f yf r yr MV C C C R (10) The following five parameters are unknown: 1f , 2f , r , 1 , 2 , r 1 2 1 2 1 1 1 1 2 2 1 0 1 0 1 0 1 0 1 1 0 0 0 0 1 0 0 0 0 0 yf yf yr f yf yf yf f rr C C C a b b cC C C s a b c R c R MV R (11) Vehicle is rotating around a point i.e. turning center CT. Forces perpendicular to wheels intersect at center of forces CF. Lines showing perpendicular tire forces from 1st ,2nd,3rd,4th,5th,6th axle are concurrent at point CF. For simplicity left and right of wheels of 1st are considered as one wheel. Similarly left and right wheels of 2nd, 3rd, 4th, 5th and 6th axle are considered as one wheel. Figure 2 shows six axle vehicle. Figure 3 shows 1st and 2nd axle wheels are steering clockwise while 5th and 6th axle wheels are steering in the opposite direction i.e. in anti-clockwise direction. 3rd and 4th axle wheels are non-steered wheels.1st, 2nd,5th , 6th axle wheels are steering as well as slipping while 3rd and 4th axle wheels are just slipping. In developing mathematical model, steering angle of 3rd and 4th axle relative to the datum line has been calculated, datum line is the reference angle to calculate remaining steering angles. Figure 3 shows two lines from CT and between 3rd and 4th axle wheels. One line is the datum line which is from CG to CT while second line is the relative line of 3rd and 4th axle which is from middle of 3rd and 4th axle and CF. The steer angle of 3rd and 4th axles reference to datum line is the angle between datum line and center line of 3rd and 4th axles. Steer angle of 3rd and 4th axles with reference to datum line is as follows: Wheels angle with reference to datum rear = 3 1( )f r c c d R R (12) The first axle angle with reference to datum can be calculated as: First axle angle with reference to datum is = 1 ( ) f a b c R (13) 4 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003016_j.compstruct.2018.11.095-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003016_j.compstruct.2018.11.095-Figure5-1.png", + "caption": "Fig. 5. Schematic view of two multiple tape arrangements. (a) Description of elements for multiple tape calculation (b) Local deviations at the start start and for maximum deviation dev,max for each tape.", + "texts": [ + " In addition, our application mechanism enables an increase of tape width. With the new application mechanism, tapes trajectories can not be continuously placed in parallel on double-curved surfaces. Due to resulting gaps and overlaps, there is a need for a rearranged multiple tape design. In the following, we describe the multiple tape optimisation problem within the specific requirements and constraints of our use case. Subsequently, an application-driven fitness criterion is defined and implemented as a two-phase iterated local search. Fig. 5 displays two exemplary multiple tape designs to describe the constraints and parameters of the optimisation. A multiple tape design for a surface (1) has to fulfill three constraints: \u2022 The surface (1) must be completely covered (However, tapes can start and end on an enlarged surface (7) for calculation reasons) \u2022 The tape trajectories must be generated with the geometry induced mechanism (equal edge lengths, see Fig. 3(b)) \u2022 A continuous overlap distance orthogonal to the application direction has to be ensured between aligned tapes (2)", + " In the experimental investigation, we determined that a maximum tape width =W 800 mmT was sufficient for the occurring Gaussian curvatures in our use case. Because of this, we fixed the tape width at =W 800 mmT . Likewise, the tape length =L 2800 mmT was fixed to ensure a traverse of the entire surface. The number of tapes NT is a result of the calculation and optimisation. In summary, we narrowed down the optimisation parameters to achieve an application-oriented multiple tape design. The start orientation start T, 1 (see Fig. 5(b)) of the first tape is the only manipulated parameter. Further tapes are aligned and subsequently slightly rotated so that all constraints, especially overlap is ensured. We aim a minimal deviation dev mean, (see Fig. 5(b)) with respect to the yz-plane for all tapes. We implemented the optimisation as a two-phase iterated local search. This implementation is a non-deterministic approach, as the geometry-induced tapes cannot be explicitly calculated. First we describe the procedure of the local search algorithm. Subsequently, we describe the trajectory generation of the aligned tapes and the calculation of the fitness criterion in detail. The algorithm is a two-phase local search, where the start orientation start T, 1 of the first tape is varied", + " In the first phase start T, 1 is rotated with a linear rough increment phase1 for a desired number of steps nphase1. For every step the aligned tapes are generated and the fitness criterion for all tapes dev mean step, , is calculated. The second phase is a refined search around the best fitness criterion ( dev mean best, , ) of the first phase. Therefore, the second phase is analogue to the first but uses a refined increment phase2 and determine the solutions between the previous and following iteration step. Fig. 5(b) schematically displays the multiple tape generation with the essential values. Fig. 6 shows the flow chart of one algorithm phase. After the first tape (T1) is generated with the initial parameters, the second tape (T2) is generated. The start point location of the second tape is shifted byWT orthogonal to the start direction from the first tape start point. The start orientation for T2 start T, 2 is initially set to start T, 1. The orientation start T, 2 has to be adjusted to guarantee minimal overlap between T1 and T2", + " This calculation is repeated until a solution > >lap lap lapmax min T T min, 1, 2 is generated. The procedure for T1 and T2 is analogously repeated for the following tapes (Ti). In cases where the tape trajectory leaves the surface or no valid solution lapmin T T, 1, 2 is determined, a trajectory generation is not possible and the solution is rejected in the following evaluation. The fitness criterion dev mean step, , is calculated using the maximum deviations dev max Ti step, , , of each single tape (Ti)(see Fig. 5(b)). The fitness criterion = =dev mean step N i N dev max Ti, , 1 1 , ,T T is calculated as the arithmetic mean of all tapes dev max T Ti step, , 1 , . Since the geometry-induced trajectory algorithm is implemented in Matlab, the two-phase iterated local search is also implemented in Matlab. The iterated local search is able to find a local minimum (within the resolution of the refined increment), in the preset limits. For each iteration step the fitness criterion dev mean step, , is exported. In addition the greatest determined overlap is exported for discussion of the results as the maximum overlap distance lapmax " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003503_cac.2018.8623527-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003503_cac.2018.8623527-Figure1-1.png", + "caption": "Fig. 1. 2-dimensional engagement geometry", + "texts": [ + " III proposed the robust IGC algorithm by employing the back-stepping control technique combining with sliding-mode method. Meanwhile, a nonlinear disturbance observer is also introduced in this section. Sec. IV shows the effectiveness of the introduced IGC algorithm proved by the mathematical simulations. And Sec. V concludes the whole paper. II. PROBLEM FORMULATION With the assumption that the missile is a point mass moving in the vertical plane with a constant velocity, the schematic diagram of planar homing engagement geometry is painted in Fig. 1. \u03b8 and V are the missile flight-angle and missile constant velocity; R and q\u03b3 are missile-target relative range and the line of sight (LOS) angle. The planar missile-target relative motion dynamics can be written as equations below ( ) ( ) cos sin R V q Rq V q \u03b3 \u03b3 \u03b3 \u03b8 \u03b8 = \u2212 \u2212 = \u2212 (1) Differentiating (1) with respect to time, one can get 2V R Rq q V R R\u03b3 \u03b3 \u03b8= \u2212 + (2) According to the missile flight principle, missile flight dynamic equations in longitudinal plane can be expressed as sin cos z z z z mV P Y mg J w M w \u03b8 \u03b1 \u03b8 \u03d1 \u03b1 \u03d1 \u03b8 = + \u2212 = = = \u2212 (3) where m denotes the missile mass, P and Y denote the thrust and aerodynamic lift, respectively, zJ and zM are the inertial moment and pitch moment, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001699_1077546318767559-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001699_1077546318767559-Figure1-1.png", + "caption": "Figure 1. Scaled structure of the rotor system in an integrally geared centrifugal compressor.", + "texts": [ + " Based on the basic parameters of a geared rotor system test rig, the meshing force and rotor vibration displacement characteristics under different unbalance parameters, loads, backlashes, and transmission errors together with the vibration of the box are obtained by the finite element method, and are verified by experiment results. The results show that the acceleration of the box can reflect the meshing state of gear pairs well, which will be helpful in the vibration control and condition monitoring of the equipment. The structure of the geared rotor system in an integrally geared compressor is shown in Figure 1, which was obtained through the similarity design method. The geared rotor system is composed of five parallel shafts which are meshed by a helical gear including one input shaft, one intermediate gear shaft, and three high speed output shafts. Both ends of the five shafts are respectively supported by an oil film bearing. The dynamic model of the five shafts geared rotor system is shown in Figure 2. The gears on I, M, O1, O2, and O3 are numbered 1\u20135 with angular velocities !1 to !5, respectively, where time-varying mesh stiffness kij\u00f0t\u00de, backlash 2bij, and transmission error eij\u00f0t\u00de are included" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001725_icit.2018.8352179-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001725_icit.2018.8352179-Figure6-1.png", + "caption": "Fig. 6. Measuring equipment", + "texts": [ + " In order to estimate the relative position of the end points of the tube, it is required to know the tensile force of the tube. Furthermore, in order to obtain the catenary number a using equation (9), the tensile force Th is required. So as to obtain the relative positional of the end points of the tube, the tensile force Ti and its angle i are required. Therefore, in this research, we measured the vertical tensile force Tiv of the tube using the load cell, and measured the angle i of the tensile force by fixing the tube to the tip of the joystick. The measuring equipment is shown in Fig. 6. As shown in Fig. 6, the joystick was fixed to one end of the load cell using fixture for joystick. A fixture for load cell was attached to the other end of the load cell. Fixtures were made by 3D printer. In order to measure the tensile force of the tube, the payload of the drone was assumed to be 1000g and a load cell that its measurement range of 5 kg was used. The total weight of measuring equipment is 82g, thus the drone can carry. The distance x [m] between the end points and the altitude difference z [m] of the tube are calculated from the measured vertical tensile force Tiv and angle i using this measuring equipment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure19-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure19-1.png", + "caption": "FIG. 19: HELICAL GEOMETRY.", + "texts": [ + " The unrolled geometry for a 3-lobed cam is shown in Fig. 18. 8 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use This procedure would work for bosses and recessed features appended on a rotary base via the creation of unrolled / re-rolled edge geometry from axial slices. Appended features more aligned axially need another tool path solution approach. For the spiral geometry shown in Fig. 19, an unroll solution approach would not be ideal as the numerous linear tool paths would have many stops and starts, and rapid moves. Depending on the AM system (i.e., a direct energy deposition system with a powder feed), this could lead to much raw material waste. Unique AM rotary, 4 axis and 5 axis tool path solutions need to be developed. This is an active area of research for planar toolpaths [22, 23], and should be extended to the rotary, and general surface domains. To effectively emulate spray coating, new bead cross section shapes should be introduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003420_oceans.2018.8604727-Figure3-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003420_oceans.2018.8604727-Figure3-1.png", + "caption": "Fig. 3. The AR2D2 AUV, with the body fixed frame (Ob, xb, yb, zb), and the earth-fixed frame (OI , xI , yI , zI).", + "texts": [ + " The Raspberry pi processes the information from the sensors, computes the control laws and sends signals to the actuators (DC motors) controlled by Pulse Width Modulation through motor drivers (Robbe Rookie 25A ESC). Real time communication is provided by Ethernet LAN interface. In Figure (2) is depicted a schematic view summarizing of the various components of the vehicle\u2019s hardware and their interactions. B. Designing and Movement Description It has designed and built the AR2D2 micro submarine, which is shown in Figure 3, with its body fixed frame (Ob, xb, yb, zb). The center (Ob) of this frame corresponds to the center of vehicle\u2019s gravity, and its axes are aligned with the main axes of vehicle\u2019s symmetry. The movement in horizontal plane is referred as surge (along xb axis) and sway (along yb axis), while heave represents the vertical motion (along zb axis). Roll, pitch, and yaw, denoted (\u03c6, \u03b8, \u03c8), are the Euler angles describing the orientation of the vehicle\u2019s body fixed frame with respect to the earth-fixed frame (OI , xI , yI , zI), while (x, y, z) denote the coordinates of the body-fixed frame center in the earth fixed frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002432_978-3-319-99522-9_8-Figure8.16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002432_978-3-319-99522-9_8-Figure8.16-1.png", + "caption": "Fig. 8.16 General layout of a 3-RRR planar parallel robot", + "texts": [ + " of the 3-RRR Planar Parallel Robot The mechanism of this robot consists of a traveling platform which is connected to the base platform by three planar kinematical chains. Each chain consists of two links attached using three revolute joints. One end of every chain is connected to the actuator, while the other is connected to the moving platform. The link connected to the actuator is called a crank and the link connecting the crank and the platform is called a coupler. Connected to all the chains, the moving platform is as an equilateral triangle and permanently carries the end-effectors. Pursuing the general schematics of a 3-RRR planar parallel manipulator (Fig. 8.16), though any three of the nine joints can be actuated for desired manipulation and the actuators under consideration, they have wisely been placed at the base of each link, in order to lighten the weight of the moving components. Based on the principle of virtual work, a recursive procedure is developed to obtain compact results concerning the time-history evolution for the input torques of the three actuators, but also for external and internal forces in the joints. Having a closed-loop structure, the planar parallel robot 3-RRR is a symmetrical mechanism composed of three planar chains with identical topology, all connecting the fixed base to the moving platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000280_j.jappmathmech.2010.01.002-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000280_j.jappmathmech.2010.01.002-Figure2-1.png", + "caption": "Fig. 2.", + "texts": [ + ", if the direction of the angular velocity vector is not the same as the direction of the axis of the disk, but the angle between the axis of the disk and the angular momentum vector is small, this angle will remain small during the entire time of motion as a consequence of the properties of regular precession. Hence, in this case, too, knowing the direction of the axis of the disk, we can determine the fixed direction in inertial space with obvious reservations. Now suppose the ideal constraints applied to the disk are realized using gimbals (see Fig. 2). We will assume that a weightless rod passes through the centre of the disk perpendicular to its plane and specifies the axis of rotation of the disk O relative to the inner gimbal. Suppose the axis of the outer gimbal of the suspension is vertical, the axis of the inner gimbal is horizontal, and the centres of gravity of the disk and each of the gimbals coincide with the centre of the suspension. We will use , and to denote the angles of rotation of the outer gimbal relative to the fixed trihedron, the inner gimbal relative to the outer gimbal, and the disk relative to the inner gimbal, respectively. We will use the letter O to denote the centre of the suspension, and we will bind the movable trihedron Oxyz to the outer gimbal (Fig. 2). We introduce the movable trihedron O , which is bound to the inner gimbal in such a manner that the following relations hold for its unit vectors: ex = e , (ey, e ) = cos . We will assume that the outer and inner gimbals are identical and that the following relations hold for the moments of inertia (about the principal axes) of the outer gimbal (Jx, Jy, Jz), the inner gimbal (J , J , J ), and the disk (A, B, C) We write the expressions for the angular velocities 1, 2 and 3 of the outer gimbal, the inner gimbal and the disk, respectively, Accordingly, the kinetic energy of each of these bodies has the form and the kinetic energy of the system is The structure of the Lagrangian, which is the same as the kinetic energy in the problem under consideration, enables us to write the three first integrals (the energy integral and two cyclic integrals): (2) The equations of motion of the system specified by these first integrals have the solution for any 0, 0, 0 and 0, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000348_12.828134-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000348_12.828134-Figure7-1.png", + "caption": "Fig. 7 (a) Illustration of the FEOET integration with a PDMS microwell-patterned substrate. (b) Experimental demonstration of an optical delivery of a 1-\u00b5L water droplet to merge with one of the target wells for dilution process. Before immerging into the oil chamber, the 5\u00d72 microwells are preloaded with Trypan blue dye.", + "texts": [ + " When the two droplets are close enough, dipole-dipole interaction causes electrocoalescence that merges these two droplets into one [19, 20]. The merged droplet is then transported on the FEOET surface to enhance mixing using droplet internal flow induced by shear stress coming from the bottom surface during transportation (see Fig. 6(e)). One unique characteristic of FEOET is that the light induced electric field pattern can penetrate into the oil layer up to hundreds of micrometers. This advantage allows light-induced DEP manipulation of droplets on top of electrically nonconductive substrates positioned on FEOET. Fig. 7 shows an example of FEOET integration with a 5\u00d72 array of PDMS microwells positioned on top of a cover glass substrate. Each microwell is 1 mm in diameter and 150 \u00b5m in depth. The Proc. of SPIE Vol. 7400 74000U-7 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 03/18/2016 Terms of Use: http://spiedigitallibrary.org/ss/TermsOfUse.aspx wells are filled with Trypan blue dye before placement on the FEOET platform. A 1-\u00b5L water droplet is optically transported on top of this PDMS microwell substrate and directed to merge with one of the target wells for dilution as shown in Fig. 7(b). The media preloaded in the wells, or carried by the droplet, can be extended to other chemicals, drugs, living cells (both adherent and non-adherent), and other biochemical reagents for wider applications. After experiments, these external substrates can be removed from the FEOET platform and replaced by a new disposable substrate to reduce cross contamination. The motion of a shadow bar-shaped image pattern allows continuous and dynamic electrowetting modulation on a LOEW surface. Fig. 8(a) shows snapshots demonstrating optical transportation of a 50-\u00b5L water droplet at a speed as fast as 17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002171_978-3-319-49574-3_12-Figure12.38-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002171_978-3-319-49574-3_12-Figure12.38-1.png", + "caption": "Fig. 12.38 Gas compression (GC)", + "texts": [ + " Gas compression cables are usually of three-core design with all three cores being located inside a steel pipe The cable is kept under a positive gas pressure (12 bar) using a pumping system. The gas pressure must be monitored using pressure gauges and alarms. This design of cable has been used successfully from the 1930s but has the disadvantage of using gas and the need to monitor the gas. These cables were mainly installed n Europe, and installation of this design for new circuits ceased in the 1970s. These cables are usually three-core in a single steel pipe. The current voltage limit is 132 kV. A typical cable cross-section is shown above in Sect. 12.8.2, Fig. 12.38. In this case, the conductor is insulated with cross-linked polyethylene (XLPE). This design of cable has been used successfully from the 1970s and is now the preferred design with most utilities as the design is relatively simple and there is no oil or gas to monitor. Depending on the voltage and the rating, these cables may be three-core or single-core cables. The current voltage limit is 600 kV a.c. A typical cable crosssection is shown above in Sect. 12.8.2, Fig. 12.39. This design of cable has been developed for use in the last few years" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003955_kem.801.65-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003955_kem.801.65-Figure1-1.png", + "caption": "Fig. 1 The Euler angles of the short fiber", + "texts": [ + " The effective elastic modulus L is an explicit function which is easy to obtain. For the case of perfectly aligned fiber with different orientated angles, a relationship between the elastic tensor \ud835\udc73\ud835\udc73ijkl of the global coordinates and \ud835\udc73\ud835\udc73i\u2032j\u2032k\u2032l\u2032 of the local coordinates was obtained: \ud835\udc3f\ud835\udc3fijkl = \u03b1\ud835\udc56\ud835\udc56\u2032\ud835\udc56\ud835\udc56\u03b1\ud835\udc57\ud835\udc57\u2032\ud835\udc57\ud835\udc57\u03b1\ud835\udc58\ud835\udc58\u2032\ud835\udc58\ud835\udc58\u03b1\ud835\udc59\ud835\udc59\u2032\ud835\udc59\ud835\udc59\ud835\udc3f\ud835\udc3fi\u2032j\u2032k\u2032l\u2032 where the transformation matrix \ud835\udf36\ud835\udf36\ud835\udc57\ud835\udc57\ud835\udc56\ud835\udc56 is as follows: \ud835\udf36\ud835\udf36\ud835\udc57\ud835\udc57\ud835\udc56\ud835\udc56 = \ufffd cos\ud835\udf11\ud835\udf11 sin\ud835\udf03\ud835\udf03cos\ud835\udf11\ud835\udf11 sin\ud835\udf03\ud835\udf03sin\ud835\udf11\ud835\udf11 \u2212sin\ud835\udf03\ud835\udf03 cos\ud835\udf03\ud835\udf03cos\ud835\udf11\ud835\udf11 cos\ud835\udf03\ud835\udf03sin\ud835\udf11\ud835\udf11 0 \u2212sin\ud835\udf11\ud835\udf11 cos\ud835\udf11\ud835\udf11 \ufffd Among the elements in the \ud835\udf36\ud835\udf36\ud835\udc57\ud835\udc57\ud835\udc56\ud835\udc56, \ud835\udf03\ud835\udf03 and \ud835\udf11\ud835\udf11 are the Euler angles as shown in Fig. 1 Finite element modelling. RVE was established as a cube and the model edge sizes of L1\u00d7L2\u00d7L3 were shown in Fig. 2. Random sequential adsorption(RSA) schemes was employed for fiber generation. The fiber was generated as a cylinder. The axis of the cylinder was employed to be the baseline in the algorithm. The start point of the middle point in the cylinder axis located in the origin of model. Three random numbers were randomly generated to represent the moving distance of the baseline along the x direction, the y direction and the z direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002907_2018-32-0052-Figure16-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002907_2018-32-0052-Figure16-1.png", + "caption": "FIGURE 16 Example of a figure and figure caption.", + "texts": [ + " The simulation used the finite element method (FEM) model to the frame, swing arm, and some part of the steering system. The engine and other heavy loads were assumed to be\u00a0rigid bodies and placed at their center of gravity positions and coupled by springs. Rotary joints were used to couple the steering system to the frame head pipe, the vehicle frame to the swing arm pivot, and front and rear axles. Regarding the coupling between the vehicle and the rider, the same method as that used in the simulation of the excitation experiment on the test bench was used. Figure 16 shows an external view of the model. \u00a9 2018 SAE International and \u00a9 2018 SAE Japan. All Rights Reserved. The FEM tire model provided by DYNA as shown in Figure 17 was used. This model consists of several representative elements. The material properties of each element were calculated by the numerical formula of the reference [8] based on the structure of the actual tire. Table 5 and Figure 18 show the characteristics of the tire itself obtained from the FEM tire model. Camber stiffness shows a value about 15% higher FIGURE 15 Comparison of test and simulation TABLE 4 Specifications of simulation model Wheelbase (mm) 1575 Caster angle (deg) 27" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000069_tee.20444-Figure14-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000069_tee.20444-Figure14-1.png", + "caption": "Fig. 14 Multipurpose fuzzy evaluation", + "texts": [ + " For each candidate ri in T\u0303n, the control instruction Cri is calculated by the cascade fuzzy control mechanism as shown in Fig. 9. And the future pose (xt+1, yt+1, \u03b8 t+1) of vehicle is predicted for each instruction candidate Cri by the kinematics model (4). Then multipurpose fuzzy evaluation is conducted for angle deflection, distance deflection and the minimal distance to obstacles. It calculates the evaluation values of all candidates and makes decision to select the one with the highest evaluation value as the control target by the following equation (Fig. 14). \u00b5D\u0303 = \u00b5T\u0303n (ri) \u2227 \u00b5angle(\u03b8) \u2227 \u00b5dist(x, y) \u2227 \u00b5obs(\u03b4) (6) The evaluation value of the operation instruction candidate which results moving in the opposite direction is reduced a half to make the vehicle select others\u2019 target candidates with higher evaluation to avoid trapping into dead loop of local minima. 4. Simulation Results In order to confirm the validity of the constructed control system based on soft target, we carried out four kinds of simulation: without any obstacle, with static obstacle, with moving obstacle and with static and moving obstacles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure10-1.png", + "caption": "Fig. 10. Motion type of 2R1T with submanifold T (z) \u00b7U(q,x,y)", + "texts": [ + " There have three rotational joints fixed on moving platform, which are locked when metamorphic parallel mechanisms with motion type T2(z) \u00b7 R(p,z) is obtained in Fig. 9. In this configuration, the moving platform has mobility three with two translations along the X and Y axes and one rotation along Z axis. Three rotational joints unlocked with moving platform and rotate link of each limb, then metamorphic mechanisms have the configuration T2(u1) \u00b7U(q1,u1,v1) \u22c2 T (3) \u00b7 U(q2,u2,v2) \u22c2 T2(v3) \u00b7 S(q3). Then keeping three rotational joints are locked when metamorphic parallel mechanisms with motion type T (z) \u00b7U(q,x,y) is obtained in Fig. 10. In this configuration, the moving platform has mobility three with one translation along the Z axis and two rotations along X and Y axes. This paper presented a new method to synthesize meta- morphic joint with different phases that was used to construct a patent metamorphic parallel mechanism with motion type of 1R2T and 2R1T. The key property of the metamorphic joint was studied through the common composing form of limbs but different mode of motion. Therefore, metamorphic joint was synthesized that can result two phases of different relative positions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003810_978-981-13-5799-2_11-Figure11.31-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003810_978-981-13-5799-2_11-Figure11.31-1.png", + "caption": "Fig. 11.31 Cornering properties and friction ellipse [35]", + "texts": [ + " When the cornering force reaches the maximum Fy max at slip angle amax, Fy max is given by Fmax y \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi lFz\u00f0 \u00de2 F2 x q : \u00f011:108\u00de Assume that the cornering force decreases at the same rate for all slip angles as shown in the left figure of Fig. 11.30. When the fore\u2013aft force Fx is applied to a tire, Eq. (11.108) can be rewritten as Fy Fy0 \u00bc A A0 \u00bc Fmax y lFz \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi lFz\u00f0 \u00de2 F2 x q lFz ; \u00f011:109\u00de Furthermore, Eq. (11.109) can be rewritten as Fx lFz 2 \u00fe Fy Fy0 2 \u00bc 1: \u00f011:110\u00de The relationship between the cornering force Fy and fore\u2013aft force Fx at slip angle a0 can be expressed by the ellipse shown in Fig. 11.31a, which is called the friction ellipse. At the slip angle amax where the cornering force is a maximum, the friction ellipse becomes the friction circle. Figure 11.31b shows that cornering properties with the combined slip can be estimated from only data of the friction circle. Comparing Fig. 11.31b with the measurements in Fig. 11.29, the cornering properties can be expressed by the friction ellipse except in the region of a high slip ratio. The drawback of the friction ellipse is that it cannot express the cornering properties at a high slip ratio. 11.4.2 Sakai\u2019s Model (1) Fundamental equations Sakai [2] developed a tire model for cornering properties under the combined slip condition with a large slip angle and large slip ratio by applying the same revisions used in Sect. 11.2. Figure 11" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000824_2009-01-0610-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000824_2009-01-0610-Figure1-1.png", + "caption": "Figure 1: Bearing geometry and degrees of freedom", + "texts": [ + " A complete spindle-bearing system is considered and analyzed using both detailed approach and mean contact angle approach. The use of this simplified model in a certain speed range is justified based on comparison between the results of the presented model and those of the full model. GEOMETRY - The model considers five degrees of freedom for the relative displacement of inner and outer races of an angular contact ball bearing. Since the degrees of freedom express the relative motion between the rings, it is convenient to assume that the outer ring is fixed. Figure 1 shows an angular contact ball bearing, its geometrical parameters and its degrees of freedom: three translations which are displacements in axial and radial directions, and two rotations about axes . Figure 2 represents forces acting on the bearing. Fi and Fo are the forces acting on the inner and outer raceway contacts, respectively. Fc is the centrifugal force which is responsible for difference between the contact angles. In addition, in rotation about the bearing axis, the balls will undergo a rolling motion as illustrated in Figure 1. These non-collinear rotations result in a gyroscopic moment [3] that causes the balls to rotate around an axis perpendicular to the plane of figure 2. The gyroscopic moment and the resulting rotation are resisted by friction. The detailed formulation of these high speed loadings, i.e. centrifugal Force and gyroscopic moment, is given in [3] and the results are used here to establish force equilibrium. Figure 2: Forces on the bearing SAE Int. J. Passeng. Cars - Mech. Syst. | Volume 2 | Issue 1 781 The internal configuration of the bearing and its components i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001904_012033-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001904_012033-Figure2-1.png", + "caption": "Figure 2. Available cross-sections", + "texts": [ + " In order to design taking account of ergonomic effect or human factors simultaneously with optimum structural strength, a walking tractor handlebar should have maximised natural frequencies so as to avoid vibration resonance from external excitation that may transmit to a user. Besides, minimising bending and torsion deflection can, to some extent, increase structural reliability. The minimisation of structural mass or volume will affect structural cost. The design variables include topology of side bar stiffeners, shape, and sizes of all bars. Fig. 1 and Fig. 2 show an initial handlebar structure which comprises the main handlebars and stiffeners. The shape of the structure is defined by the parameters; y1, y2, y3, y4, y5, z1 and z2 while sizing of the structure is defined by the parameter; r1, r2, r3, r4, r5, r6, r7, r8, r9, r10 t1, t2, t3 and t4. The parameters A, B, C, D, E, F and G [0, 1] are used to define shape of the bars, which can be either a round or a square tubing cross-sections as presented in Table 1, while the parameters H, I, and J are also used to define shape of bars, which can be either a round or a square bar cross-sections as presented Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001529_s12008-018-0471-y-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001529_s12008-018-0471-y-Figure1-1.png", + "caption": "Fig. 1 Self-lubricating joint bearing", + "texts": [ + "com 1 School of Mechatronic Engineering and Automation, Shanghai University, Shanghai 200072, China 2 Shanghai Bearing Technology Research Institute, Shanghai 201800, China tion (MEMS) wear simulation. The experimental results and the simulation results are very consistent [5]. Bortoleto used the finite element software to write user subroutines and used Archardwearmodel to simulate thematerial removal process of the pin on disc [6]. Due to high carrying capacity and long service life, selflubricating joint bearing is widely used in aerospace and other fields. Self-lubricating joint bearing is shown in Fig. 1, including the inner ring, fabric liner and outer ring. In the process of design and production of products, the life of self-lubricating joint bearings is key factor. The main failure form of joint bearing is wear failure, and its effective life depends on the predetermined limit of wear depth of the self-lubricating liner. In order to judge the effective life of a self-lubricating joint bearing product, the wear experiment under a predetermined standard working condition is required [7]. If thewear depth of self-lubricating joint bearing liner meets the design specification, the product is qualified [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003949_iscid.2018.10152-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003949_iscid.2018.10152-Figure5-1.png", + "caption": "Figure 5. Equivalent strains nephogram", + "texts": [ + " Higher stress is concentrated in the hub and the reinforcement ring, especially the maximum stress occurs just below the load of the inner ring of the reinforcement ring, which is also the monitoring point RF. Therefore, it is necessary to check the strength of the reinforcing ring. The reinforcing ring material is Q235, and its yield strength is 235 MPa. According to the safety requirements, the strength of the driving drum must be less than 70% of the material distinguishing strength, which is 164.5 MPa. max=112.8MPa< [ ]=164.5MPa Therefore, the driving drum meets the strength requirements. Figure 5 (a) is an equivalent strain nephogram of the driving drum under 500 times the load. For convenience of observation, the drum is cut symmetrically along the XOY plane to show the strain variation inside the drum, as shown in Figure 5 (b). It can be clearly seen from Figure 5 that the minimum strain of the roller is 0.0049 mm and the maximum value is 0.489 mm. When the strain is large, it is concentrated in the lower half of the cylinder, which is also far away from the load, especially the maximum stress appears in the lower half of the cylinder on both sides of the reinforcement ring. Therefore, it is necessary to check the deflection of the cylinder. According to safety requirements, the deflection of the transmission cylinder must be less than [ ]. max= 0.489 mm< [ ] = L/2500=0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure10-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure10-1.png", + "caption": "FIG. 10: (A) SOLID 90\u00b0 TORUS, (B) TWO SECTION", + "texts": [ + " The slices for each segment in Fig. 9 (c) are parallel to each build plane. The amount of support material for this 90\u00b0 torus would be minimal if it were built \u2018laying down\u2019 or the top view orientation, but if the torus were a thin walled component, or had some internal geometry, support material would be required in any build orientation, and it would be challenging to remove. Employing a rotary positioning system, and including side tilt, provides a fabrication solution for the geometry shown in Fig. 10 (b) while using conventional slicing and tool path options. This technique can be employed for non-symmetric features. For example, the flange for the thin shelled bullet component can also be fabricated using this approach, building up regions onto an inner ring (Fig. 11). Four segments, with a maximum 45\u00b0 overhang, are determined, and the segments built up with planar slicing. 6 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002242_978-3-319-99262-4_29-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002242_978-3-319-99262-4_29-Figure1-1.png", + "caption": "Fig. 1. Rotor seal model", + "texts": [ + " The fundamentals of rotordynamics, modeling, and simulation are well described in [18\u201320]. For the sake of clarity, we write the formulation and the analysis of the rotor seal system here again, which is already presented in [21]. In the following, the Jeffcott/Laval rotor model is used to described the dynamic behavior of the system. Like the investigations of Black [2], Childs [20], and Muszynska [19], the seals are modeled using rotordynamic seal coefficients. The simplified rotor seal model, see Fig. 1, consists of a linear elastic, massless shaft with a mass disk supported by two rigid bearings. The seals, represented by their rotordynamic coefficients (mass, stiffness, and damping) are directly connected to the shaft. Using the rotor displacement [x y]T and projecting all forces on the disk, the equation of motion is as follows:[ mr + mxx 0 0 mr + myy ] [ x\u0308 y\u0308 ] + [ cxx cxy cyx cyy ] [ x\u0307 y\u0307 ] + [ kr + kxx kxy kyx kr + kyy ] [ x y ] = h (1) where mxx, cxx, and kxx are the seal coefficients of direct mass, damping, and stiffness of the seals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003738_978-3-030-04975-1_7-Figure5-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003738_978-3-030-04975-1_7-Figure5-1.png", + "caption": "Fig. 5. Distribution of intertooth forces Fi", + "texts": [ + " It is also possible to narrow down the search field for the value of that clearance by selecting transmission elements, e.g. a roller featuring the same deviation and a drive shaft featuring the same deviation of the eccentric. During the operation of the cycloidal gear transmission, MK torque is generated and it acts on the planetary gear. This torque generates intertooth forces Fi which occur between the radius g rollers and the hypocycloidal curve that gets in contact with them. The distribution of forces is shown in Fig. 5. The load (torque MK) is transferred by half of the tooth pairs on the active side of the backlash-free gear transmission. In the case of the gear transmission featuring intertooth clearances \u0394i (backlash), the number of collaborating tooth pairs decreases depending on the size of that clearance. Each intertooth force Fi is a linear displacement function di proportional to force Fmax, which can be expressed as: Fi \u00bc Fmax hi hmax \u00bc Fmax di Di dmax \u00f012\u00de where: hi \u2013 rolling radius, hmax \u2013 max rolling radius, di \u2013 displacement, dmax \u2013maximum contact displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001121_2008-01-1408-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001121_2008-01-1408-Figure1-1.png", + "caption": "Figure 1: Acting forces and moments to a cambered tire", + "texts": [], + "surrounding_texts": [ + "By mandate of the Engineering Meetings Board, this paper has been approved for SAE publication upon completion of a peer review process by a minimum of three (3) industry experts under the supervision of the session organizer.\nAll rights reserved. No part of this publication may be reproduced, stored in a retrieval system, or transmitted, in any form or by any means, electronic, mechanical, photocopying, recording, or otherwise, without the prior written permission of SAE.\nFor permission and licensing requests contact:\nSAE Permissions 400 Commonwealth Drive Warrendale, PA 15096-0001-USA Email: permissions@sae.org Tel: 724-772-4028 Fax: 724-776-3036\nFor multiple print copies contact:\nSAE Customer Service Tel: 877-606-7323 (inside USA and Canada) Tel: 724-776-4970 (outside USA) Fax: 724-776-0790 Email: CustomerService@sae.org\nISSN 0148-7191 Copyright \u00a9 2008 SAE International Positions and opinions advanced in this paper are those of the author(s) and not necessarily those of SAE. The author is solely responsible for the content of the paper. A process is available by which discussions will be printed with the paper if it is published in SAE Transactions.\nPersons wishing to submit papers to be considered for presentation or publication by SAE should send the manuscript or a 300 word abstract of a proposed manuscript to: Secretary, Engineering Meetings Board, SAE.\nPrinted in USA", + "Copyright \u00a9 2008 SAE International\nABSTRACT\nBased on the simulation results of tire rolling over perpendicular cleats by MPTM model, in present paper, a series of simulation results of tire rolling over oblique cleats with different angles are given. For that, the Modal Parameter Tire Camber property Model is established. For the appraisement of comparison between simulation and experimental results a problem concern the validation test is pointed out. In the end, simulation results of tire rolling over a series of continuous cleats are given.\nINTRODUCTION\nSeveral high level analytical models are developed to meet the growing demand in automotive virtual proving ground simulation technology. They are SWIFT-Tyre [1], RMOD-K [2], FTIRE [3] and MPTM [4-12]. In paper [12], a series of simulation results of tire rolling over perpendicular cleats by MPTM model are given. They are quite agreed with the experimental results in both time and frequency domains. And a series of simulation capabilities of the model are presented. But for full virtual proving ground simulation, a 3-dimensional model should be developed. For example, simulation of tire rolling over oblique cleats is lateral force and moment response concern. For that, the modal parameter tire camber property model is established on the basis of the vertical model. Lateral modal parameters are introduced and tread width is considered. A steady state case calculated for a tire 205/55R16 with camber angle and fixed rim center in horizontal position is given. Vertical force is roughly verified by experiment in range of vertical deformation to 0.025m. Vertical, cornering and camber model are synthesized, and a three dimensional MPTM is formed. Simulations of a tire 205/55R16 rolling over oblique cleats with different incline angles are performed. A series of calculation results, such as: vertical, longitudinal, lateral forces responses; moment responses about vertical, longitudinal axis and wheel rotation speed varying are given in both time and frequency domains. All dynamic simulation results are similar to the results in reference [13] in which effective road plane inputs are used. Because of lacking of experimental results to compare with, the\ncomparison between simulation and experimental results of the same tire rolling over perpendicular cleat of 0.02m height with 2km/h speed is given. It can be seen that the amplitudes of calculated vertical and longitudinal response forces agree with the experiments very well. But there is a problem exists in time process, which is to be analyzed and very important for test validation of simulation results. In the end, an example of simulation of tire rolling over a series of continuous cleats is performed. But our results still have to be verified by experiments.\nCAMBER PROPERTY TIRE MODEL\n2008-01-1408\nThe 3-Dimensional Modal Parameter Tire Model and Simulation of Tire Rolling Over Oblique Cleats\nChengjian Fan, Dihua Guan and Baojiang Li Tsinghua University", + "Where: fr - radial force vector; ft - tangential force vector; Hrr, Hrt - radial and tangential displacement mobility matrix by radial excitation; Htr, Htt - radial and tangential displacement mobility matrix by tangential excitation.\nAccording to the modal theory, if l is the excitation point and p is the response point, transfer function with N orders is:\n2 1 j\nN lk pk\nlp k k k k\nH K M C\n(4)\nWhere, lr and pr , rK , rm , rC are mode shape, modal stiffness, modal mass and modal damping factor respectively.\nThe tread radial hr and tangential ht deformation in the contact patch are:\nr n r\nt m t\nh f k\nh f k\n(5)\nWhere: fn- vertical force; fm- longitudinal force; kr - normal stiffness of tread; kt - shear stiffness of tread.\nIn camber model, in which lateral modal parameters of tire are taken in to account, formula (3) becomes formula (6), and the total radial r, tangential t and lateral l deformation displacements are:" + ] + }, + { + "image_filename": "designv11_92_0002357_remar.2018.8449843-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002357_remar.2018.8449843-Figure1-1.png", + "caption": "Fig. 1. Metamorphic joint", + "texts": [ + " The type synthesis of motion type T2(w1) \u00b7U(p1,w1,u1) has been discussed and one of limbs is generated through MT2(w1)\u00b7U(p1,w1,u1) = MPL(w1) \u00b7 MC(p1,u1)/T (u1), where C(p1,u1)/T (u1) is quotient group in literature[14], PL(w1) and C(p1,u1)/T (u1) are generated by R(p11,w1) \u00b7 R(p12,w1) \u00b7R(p13,w1) and R(p14,u1), respectively. Similarly, T2(u1) \u00b7U(q1,u1,v1) limb is generated by the same form MT2(u1)\u00b7U(q1,u1,v1) = MPL(u1) \u00b7MC(q1,v1)/T (v1), where PL(u1) and C(q1,v1)/T (v1) are generated by R(q11,u1) \u00b7R(q12,u1) \u00b7 R(q13,u1) and R(q14,v1), respectively. Metamorphic joint is synthesized which can realize different motion types combining with constraint conditions of T2(w1) \u00b7U(p1,w1,u1) limb and T2(u1) \u00b7U(q1,u1,v1) limb in Fig.1. This metamorphic joint is combined by link a\u223c d and three revolute joints with axis 1, axis 2 and axis 3. Link a is fixed in base or moving platform, link d is connected to other links. In the metamorphic joint, rotational axis 1, rotational axis 2 and rotational axis 3 are perpendicular to each other in the initial state. The output degree of freedom is two when revolute joint of axis 1 or axis 2 is locked. There have two phases when changing position of link b and link c through locking revolute joint of axis 1 or axis 2 in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003686_sdpc.2018.8664975-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003686_sdpc.2018.8664975-Figure2-1.png", + "caption": "Fig. 2. Planetary gear fault with a cracked tooth.", + "texts": [ + " EXPERIMENT AND RESULT In this study, the experimental vibration data of planetary gear train is generated by a helicopter transmission test rig, as shown in Figure 1. The experimental system is composed of a drive motor, a main rotor gearbox containing a planetary gear train, an adjustable load and a vibration testing system. The employed vibration data was collected using an accelerometer attached to the main rotor gearbox housing with a sampling rate of 5kHz. To demonstrate the performance of the proposed method, the vibration signal of a planetary gear with an injected crack fault is considered, as shown in Figure 2. Figure 3 shows the experiment results, including the waveforms of the raw signal, the deconvolution signal after FAO-MCKD preprocessing, the filtered signal after OGS operation, and the envelope spectrum of the filtered signal. In the raw signal waveform, (Figure 3a), a small amount of large-amplitude impact emerge without periodicity, demonstrating there is something unknown wrong with a part of the system. In the deconvolution signal waveform (Figure 3b), which is achieved by FAO-MCKD preprocessing, the large amplitude impact components originally masked by the interference and noise is effectively highlighted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001813_978-981-10-8306-8_13-Figure13.6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001813_978-981-10-8306-8_13-Figure13.6-1.png", + "caption": "Fig. 13.6 Multiple normal directions at a sharp corner of the Tresca p diagram obtained considering a smooth curve converged to a sharp corner as its limit", + "texts": [ + "35) and while considering the stress distribution shown in Fig. 12.6. Also, derive that a \u00bc 1 2 or a \u00bc 1 4, respectively, when the reference state is a simple tension (and balanced biaxial) state or a pure shear state. The Tresca p diagrams of the yield surface and the plastic strain increment surface are plotted based on the dual normality rules in Fig. 13.5. Note that there are multiple normal directions at the corners, which are obtained by considering a smooth curve that converges at each corner as its limit as shown in Fig. 13.6. Therefore, plastic strain increments are not unique for the simple tension or balanced biaxial stress states. Also, for the pure shear stress state, the plane strain deformation is obtained, for which one of three plastic strain increment components vanishes. However, there are multiple stress states for plane strain deformation from the simple tension to the balanced biaxial. For the pure shear stress state, deformation of the Tresca case and all incompressible, isotropic and symmetric yield functions including the von Mises case is the same as plane strain deformation, even though the stress state for the plane strain deformation is non-unique for the Tresca case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002829_detc2018-85860-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002829_detc2018-85860-Figure2-1.png", + "caption": "Figure 2. Schematic of the cast, the cutting lines, and the 4 final pieces to collect", + "texts": [ + " Due to the huge variety of the pieces, and the different ways of cutting, it is technically impossible build an automation system to be able to perform all the cuts, and it is not economically feasible to automate the process for all the variety of pieces with one robotic cell for each. Thus, the expected solution should provide the existing flexibility in operation, as it is provided by human operator, and should bring the assistance to the operators for the hard tasks. The cutting task is depicted in the Figure 2 in a schematic way. The whole piece comes as the input, and the operator will perform 4 cuts with the cutting tool to separate the final pieces (4 cylindrical shape) from the remaining. The difficulty of this task is the weight of the cutting tool, plus the force to apply for the cutting. In addition to this assistance, there is another objective for the robotic transformation which is the gain in productivity, measured by the average number of pieces cut in the given time. This foundry company mentioned being in the situation that the demand is higher than their production capacity, and the bottleneck in the whole process is the time consuming manual tasks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001617_s40799-018-0232-7-Figure12-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001617_s40799-018-0232-7-Figure12-1.png", + "caption": "Fig. 12 Results of the finite element model with maximum load, (a) Normal stress, (b) Tangential strain, with respect to the local cylindrical coordinate system of the solid cylinder", + "texts": [ + " However, they diverge as the load increases, giving a mean difference of approximately 27%, between numerical and direct experimental results; once the latter are transversely corrected, the mean difference is reduced to approximately 11%. On this basis, the finite element model is assumed to be validated; thus the model is reliable for assessing other results of interest such as normal to axis stress and tangential strain with respect to the local cylindrical coordinate system of the solid cylinder, as shown in Fig. 12. From the results and conclusions reported by the students, most of them agreed on the following: 1) bonding and instrumenting a T type strain gage on a circular geometry was more challenging and providedmore information than a single grid strain gage bonded to a flat surface; 2) learning or remembering Computer Aided Drawing, Finite Element Analysis and Embedded Data Acquisition software, along with an analytical method, allows them to develop critical and hands-on skills, as well as, to improve their abilities for structural analysis and design; finally, 3) numerical, analytical and experimental methods may not reach a 100% concordance on their results, but an appropriate validation allows them to be confident on their developed models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000749_roman.2009.5326259-Figure6-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000749_roman.2009.5326259-Figure6-1.png", + "caption": "Fig. 6. Single inverted pendulum", + "texts": [ + " It is possible to write the objectives as in the following: \u2022 Task 1: e\u03071 = 0 = JRq\u0307R \u2212JLq\u0307L \u2022 Task 2: e\u03072 = x\u0307r = JCq\u0307C +JRq\u0307R Writing the tasks as a conventional way, it is: e\u0307 = Jq\u0307 (3) Equation 3 can be written in matrix form: [ e\u03071 e\u03072 ] = [ 0 JR \u2212JL JC JR 0 ] q\u0307C q\u0307R q\u0307L (4) In the most general case, the J matrix in Eq. 3 is composed by two matrixes 6 by 11. This means that with these assumptions, using 6 DOFs for the closed kinematic chain, it is only possible to specify the position and two orientations for the end-effector. In a very simplified way, the dynamic model of the humanoid robot RH-1 can be considered similar to that of the inverted pendulum in Fig. 6. The similarity is established under the following assumptions. The mass of the humanoid (m) is concentrated at its mass centre (tip of the pendulum), which is at a distance l from the floor. The mass of the rigid link is then considered negligible. Besides, the action (torque T) that allows the mass m to move a specific angle \u03b8 at a speed \u03b8\u0307 (movement of the mass centre during the walking action) is effected by a servomotor (ankle of the humanoid robot) fixed at the end of the link (floor). This servomotor performs the control action to ensure the stability of the system during the walking action" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0003370_imece2018-86461-Figure11-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0003370_imece2018-86461-Figure11-1.png", + "caption": "FIG. 11: SECTIONING THE FLANGE, AND BUILDING 4 REGIONS ONTO A RING.", + "texts": [ + " The amount of support material for this 90\u00b0 torus would be minimal if it were built \u2018laying down\u2019 or the top view orientation, but if the torus were a thin walled component, or had some internal geometry, support material would be required in any build orientation, and it would be challenging to remove. Employing a rotary positioning system, and including side tilt, provides a fabrication solution for the geometry shown in Fig. 10 (b) while using conventional slicing and tool path options. This technique can be employed for non-symmetric features. For example, the flange for the thin shelled bullet component can also be fabricated using this approach, building up regions onto an inner ring (Fig. 11). Four segments, with a maximum 45\u00b0 overhang, are determined, and the segments built up with planar slicing. 6 Copyright \u00a9 2018 ASME Downloaded From: https://proceedings.asmedigitalcollection.asme.org on 02/13/2019 Terms of Use: http://www.asme.org/about-asme/terms-of-use This solution will have a longer build time due to extraneous feed and rapid moves, and there will be issues related to determining interference conditions and with blending between the regions. This segmentation choice is presented to show potential build issues with strategy. There are problems with varying heights for the top and bottom flange edges (encircled regions in Fig. 11). The number of slices for a component/subcomponent will consist of an integer \u2018ceiling\u2019 value of the feature height divided by the slice layer thickness; hence, a local built up area may result when introducing a segmentation strategy. This situation has occurred here: the number of required layers to build up the top and bottom segments resulted in a localized proud region. Although this case study generates a sub-optimal solution, it is included to illustrate the solution potentials for components with complex 3D shapes and using partitioning to reframe the problem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001086_icemi.2009.5274320-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001086_icemi.2009.5274320-Figure1-1.png", + "caption": "Fig. 1. Airframe axes", + "texts": [ + " In Section 2, the dynamic model of the missile is established, Section 3 give the frame of the missile NCS. In Section 4, design the missile NCS using time-delay switched system with packet dropout compensatory model and the simulation results are given in Section 5. Finally we conclude with Section 6. II. MISSILE MODEL In this paper, use the example of a tail-controlled tactical missile in the cruciform fin configuration in Ref. 3-348 _____________________________ 978-1-4244-3864-8/09/$25.00 \u00a92009 IEEE The mathematical description of the missile of the horizontal motion (on the xy plane in Fig. 1.) is the following nonlinear model: v r v v -1 0 y 0 y v r -1 z 0 n 0 n v = y (M, , )v -Ur + y (M, , ) 1 = m V S C v+V C -Ur 2 r = n (M, , )v+ n (M, , )r + n (M, , ) 1 1 I V Sd dC r +C v+V C n2 2 (1) where v is the sideslip velocity, r is the body rate, is the rudder fin defections angle, vy , y are semi-non-dimensional force derivative due to sideslip velocity and fin angle, vn , n and rn are semi-non-dimensional moment derivatives due to sideslip velocity, fin angle and body rate. U is the longitudinal velocity, 125m kg is the missile mass" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000452_304-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000452_304-Figure1-1.png", + "caption": "Figure 1.", + "texts": [ + " It is a generalization of the simple inflexional elastica, and includes the latter as a special case. A naturally straight elastic rod ABCDE of length 2L and uniform flexural rigidity B is deformed by two equal and opposite forces P acting at its ends A and E, and directed towards the initial position of the centre C. In addition, forces of magnitude F, acting in a direction perpendicular to the original line of the rod, are applied at B and D, which are on opposite sides of C and equidistant from it. The ends of the rod are constrained to move along its initial line. Figure 1 shows the rod in its equilibrium position, together with the system of forces acting on it. Take rectangillar axes Oxy as shown, with Ox along the initial line of the rod, and let T be any point on AC, where the arc length AT is s. Suppose the tangent at T makes an angle f? with the positive direction of Ox. Taking moments about T for the part AT of the rod we obtain = pU - qx, - xB), eB Q e G 7T, (2) where 8, is the value of 0 at A, etc., (xA, yA) are the coordinates of A, etc., and dashes denote differentiation with respect to s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001674_978-3-319-76138-1_3-Figure3.21-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001674_978-3-319-76138-1_3-Figure3.21-1.png", + "caption": "Fig. 3.21 Iterative projections computed by the Dykstra algorithm from C to S and back. Convergency towards a feasible solution in F (left). Situation outside the wrench-feasible workspace without convergency (right)", + "texts": [ + " These forces are firstly projected onto the solution space S. The resulting force distribution is projected back onto the cube C. Hassan showed that the algorithm converges to a force distribution with a minimal distance between S and C. If the sets are intersecting, this minimal distance is 0 and the desired solution is found. If no improvement on the force distribution can be made by this two-step projection, then the set F is empty and thus no solution exists. The alternating projects onto the cube C and the solution space S are illustrated in Fig. 3.21. The first projection PS from an arbitrary force f (i) onto the solution space S is achieved by PS : f (i+1) = ( I \u2212 A+TAT ) f (i) \u2212 A+TwP . (3.50) The second projection of theDykstra algorithm PC onto the cube C is simply achieved by limiting each component f j of the vector f (i) to the force limits f min and f max, respectively, as follows PC : f (i+1) = [ f\u03021, . . . , f\u0302m]T (3.51) with f\u0302 j = \u23a7 \u23a8 \u23a9 f min f j < f min f max f j > f max f j otherwise . (3.52) Hassan shows that the algorithm converges in general but practical tests and also the figures in Hassan\u2019s paper indicate that the rate of convergency is very slow and also a test implementation indicated a high number of iterations until convergency was reached" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000922_acc.2008.4586521-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000922_acc.2008.4586521-Figure1-1.png", + "caption": "Fig. 1. Single-Link Robot Arm.", + "texts": [ + " It is a well known result of the ISS theory that the origin of such a cascade system is globally asymptotically stable if the driven system is ISS [6, Lemma 4.7]. In this section the theoretical results of the previous sections are applied to design nonlinear output feedback laws for various control problems. This demonstrates that the proposed class of (inverse) optimal state feedback is not too restrictive and that it can be applied to practical control problems. A model for a single-link robot arm, shown in Figure 1, is given by [3] x\u03071 = x2 x\u03072 = x3 \u2212 F2 J2 x2 \u2212 K J2 x1 \u2212 mgd J2 (cos x1 \u2212 1) x\u03073 = x4 x\u03074 = K2 J1J2N2 x1 \u2212 K J2N x3 \u2212 F1 J1 x4 + u, (35) where x1 and x3 represent the angular at the side of the motor respectively the arm and x2 and x4 represent the appropriate angular velocities. The mass m, positioned on the arm, is moved by a electrical motor M . The parameter K is the stiffness of the spring while N , J1, J2, F1, F2 are constants representing transmission, inertia, and friction. In the following the systems parameters are chosen as K = 1, J1 = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002332_speedam.2018.8445270-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002332_speedam.2018.8445270-Figure1-1.png", + "caption": "Fig. 1. Basic motor model", + "texts": [ + " As the magnetic flux is reduced in this motor, the back electromotive force (EMF) can also be reduced in the high speed range. Accordingly, the motor can extend the speed range by controlling the magnetic force of the permanent magnet. In this paper, we have studied a variable-field type motor that can magnetize and demagnetize the low coercive force magnet by the pulse current. Finally, we propose a new motor model with structural features, and present characteristics of this motor obtained by FEA simulations. II. MOTOR MODEL Fig. 1 shows a structure of the basic motor model used for comparison in this study, and TABLE I shows specifications of the basic model. This model is the three-phase, 6-pole, 36-slot structure, and the distributed windings are employed. This motor uses Nd-Fe-B magnets, and are used for industry applications. Further Nd-Fe-B magnets are used in this model. The variable-field motor discussed in this study uses two kinds of magnets. Each magnet has different coercive force. When the motor reaches a certain speed, the pulse current is supplied to the motor, and it changes the magnetized direction of the low coercive force magnet or weakens the magnetization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000012_6.2008-4509-Figure9-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000012_6.2008-4509-Figure9-1.png", + "caption": "Figure 9- Tandem Seal Assembly", + "texts": [], + "surrounding_texts": [ + "American Institute of Aeronautics and Astronautics\n092407\n7", + "American Institute of Aeronautics and Astronautics\n092407\n8\nTable 1 shows the seal combinations tested at Rexnord Corp. Figures 31 thru 36 show the performance of several seal combinations. Testing was completed at Rexnord which resulted in the seals been tested to 12,000 rpm. The approximate seal inlet temperature was 80o F", + "American Institute of Aeronautics and Astronautics\n092407\n9\nIII. Test Data/ Results Testing was completed for the seal configurations shown in Table 1. Figure 10 shows the results of testing for all of the seals except the carbon bore seals for non-rotational conditions and different pressure levels. As shown in Figure 10, the split ring brush seals had the largest level of air flow. Based on these results, it was decided to review the data in more depth for the solid 360o seal in the single and tandem configuration and compare them to the metal labyrinth seals and carbon seals." + ] + }, + { + "image_filename": "designv11_92_0001372_peds.2009.5385681-Figure7-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001372_peds.2009.5385681-Figure7-1.png", + "caption": "Figure 7. Sensorless Brushless DC Motor.", + "texts": [ + " The data processing, command transmitting, mathematical computation, output estimation, and interface communication are integrated in the DSP responsibility. In this way, the controller has higher flexibility with complicated control algorithm. The test motor specification of the draught blower was shown in table 2. In the experimental setting, the input voltage is 156 DCV, and the rotor speed is set about 900 rpm under draught blower condition. The PWM switching frequency is 15 kHz. Testing sensorless BLDC motor and the draught blower are shown in Fig. 7 and Fig. 8 respectively. In the experiment, the driving signal closed duration at the zero-crossing point are setting 10\u00ba, 15\u00ba, and 30\u00ba electrical degree respectively. The speed command is setting from 100 to 900 rpm. Each time the command increasing 100 rpm, the system current, power consumption, phase voltage, and phase current are recorded. Closing 10\u00ba electrical degree at ZCP phase current waveforms are shown in Fig. 9 with 500 rpm and 900 rpm. The experimental results that compare with conventional six-step square driving signal without current feedback control are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002881_978-3-030-03451-1_43-Figure4-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002881_978-3-030-03451-1_43-Figure4-1.png", + "caption": "Fig. 4. Flow front development using distribution medium", + "texts": [ + " The flow front during the impregnation of UD-glass laid fabric and a longitudinal flow to the fibers is due to two effects. In the macro-pore, a pressure gradient generates a viscous flow of the resin. Due to the channels between the rovings, the local permeability is high. In the micro-pore the capillary effect dominates. This moves the resin through the roving. The local permeability is significantly lower in this micro-pore than in the macro-pore [10, 16]. The shape of the flow front depends on the vacuum contribution. With a distribution medium, the flow front develops as shown in Fig. 4. The significantly higher permeability of the distribution medium allows a higher flow rate of the resin. This results in a change in the flow direction. The resin initially flows through the distribution medium. Subsequently, it flows from the distribution medium transversely into the fiber material. It provides a three-dimensional resin flow. [3, 17, 18] Without a distribution medium, a two-dimensional resin flow sets in. The flow front runs along the fiber. For the infiltration, only the capillary effect is responsible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0001737_1.5032985-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0001737_1.5032985-Figure1-1.png", + "caption": "FIGURE 1. Application of lubricants and friction modifiers at wheel-rail contact interfaces", + "texts": [ + " This system finds its major application during Autumn season, when falling leaves from the trees get onto the rails, effectively decoupling the wheels of the trains from the rails. Traction gels also provide immunity against iron oxide formation on rails and subsequent corrosion of the tracks. At the contact of wheel flanges and rails, lubrication is generally applied for reduction of frictional stresses generated during rail curve negotiations. Lubricants also provide a smooth gliding interface for the wheels to transit through railway turnouts. However, friction modifiers are applied at the wheel tread - railhead contact, as shown in Fig. 1, and the optimum traction is always maintained. For the cases where positive friction is required, top-of-rail positive friction modifiers are used for improving the adhesion characteristics and achieving positive creep. Modifying the rail surface friction using top-of-rail friction modifiers are much easier than making amendments in the wheel-rail profiles for delivering sufficient traction. In the case of contacts with high grades of friction, lubrication is generally applied in the boundary regime, mixed lubrication regime or in the hydrodynamic regime, based on the Stribeck curve, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0002055_978-3-319-79005-3_25-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002055_978-3-319-79005-3_25-Figure1-1.png", + "caption": "Fig. 1 Principal line based and Cartesian coordinate systems", + "texts": [ + " It has been shown in [5] that it is possible to choose such a principal lines based coordinate system that hb\u03b7h\u03be 1 (4) where b (t + sin \u03d5) / (t \u2212 sin \u03d5), t 1 if \u03c3\u03be > \u03c3\u03b7 and t \u22121 if \u03c3\u03be < \u03c3\u03b7. Any orthogonal net satisfying (4) determines a net of principal stress trajectories giving a solution to the equilibrium equations and (3). It is evident that Eq. (4) reduces to Eq. (2) at \u03d5 0. Equation (4) is also valid under plane stress conditions [2]. Introduce a Cartesian coordinate system (x, y). Let \u03c8 be the orientation of the \u03be\u2212 lines relative to the x-axis, measured anticlockwise positive from the x-axis (Fig. 1). It follows from the geometry of this figure that \u2202x \u2202\u03be h\u03be cos\u03c8, \u2202y \u2202\u03be h\u03be sin\u03c8, \u2202x \u2202\u03b7 \u2212h\u03b7 sin\u03c8, \u2202y \u2202\u03b7 h\u03b7 cos\u03c8. (5) Consider the yield criterion (1). Substituting Eq. (2) into Eq. (5) yields \u2202x \u2202\u03be h cos\u03c8, \u2202y \u2202\u03be h sin\u03c8, \u2202x \u2202\u03b7 \u2212 sin\u03c8 h , \u2202y \u2202\u03b7 cos\u03c8 h . (6) Here h \u2261 h\u03be . The compatibility equations are \u22022x \u2202\u03be\u2202\u03b7 \u22022x \u2202\u03b7\u2202\u03be , \u22022y \u2202\u03be\u2202\u03b7 \u22022y \u2202\u03b7\u2202\u03be . (7) Substituting Eq. (6) into Eq. (7) results in cos\u03c8 \u2202h \u2202\u03b7 \u2212 h sin\u03c8 \u2202\u03c8 \u2202\u03b7 sin\u03c8 h2 \u2202h \u2202\u03be \u2212 cos\u03c8 h \u2202\u03c8 \u2202\u03be , sin\u03c8 \u2202h \u2202\u03b7 + h cos\u03c8 \u2202\u03c8 \u2202\u03b7 \u2212cos\u03c8 h2 \u2202h \u2202\u03be \u2212 sin\u03c8 h \u2202\u03c8 \u2202\u03be " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_92_0000978_aim.2009.5229721-Figure2-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0000978_aim.2009.5229721-Figure2-1.png", + "caption": "Fig. 2. T2R1-type parallel manipulator with decoupled and bifurcated planar-spatial motion of the moving platform: constraint singularity (a), branch with planar motion (b), branch with spatial motion of the moving platform (c), limb topology P R \u22a5 P \u22a5 R \u22a5 R-P \u22a5 R R \u22a5 R-P R RS.", + "texts": [], + "surrounding_texts": [ + "The term of constraint singularities have been recently coined [25] to characterize the configuration of lower mobility parallel manipulators in which both the connectivity of the moving platform and the mobility of the parallel mechanism increase their instantaneous values. When bifurcation occurs in a constraint singularity, the mechanism can reach different configurations, called branches, and have different independent motions of the moving platform. A branch refers to a free-of-singularity configuration of the mechanism in which each structural parameter keeps the same value. For this reason, this value is called global of full-cycle value for a branch. Two types of bifurcation have been defined in [26]. A bifurcation of type BCS1 occurs when the parallel mechanism F \u2190 G1-\u2026-Gj-\u2026-Gk, get out from a constraint singularity (CS) in different branches characterized by the same mobility and the same connectivity of the moving platform but with different bases of the vector space of relative velocities between the moving and the fixed platforms. A bifurcation of type BCS2 occurs when a parallel mechanism F \u2190 G1-\u2026-Gj-\u2026-Gk, get out from a constraint singularity in different branches characterized by distinct values of mobility and connectivity of the moving platform. The parallel manipulators presented in this paper have constraint singularities with bifurcation of type BCS1. Bifurcation of type BCS2 occurs in kinematotropic mechanisms. The finite displacements and the velocities in the actuated joints are denoted by qi and iq , and the linear and angular velocities of the characteristic point H situated on the moving platform by 1v x= , 2v y= and \u03b1\u03c9 \u03c9 \u03b1= = or \u03b4\u03c9 \u03c9 \u03b4= = . In both branches the moving platform undergoes a planar translation but the rotation axis is different in the two branches. In the first case, the rotation velocity \u03b1\u03c9 \u03c9 \u03b1= = is parallel to the x-axis and the moving platform undergoes a spatial motion. In the second case, the rotation velocity \u03b4\u03c9 \u03c9 \u03b4= = is parallel to the z-axis and the moving platform undergoes a planar motion. With these notations, the linear mapping between the actuated joint velocity space and the end-effector velocity space for the T2R1-type PMs is defined by: [ ] 1 1 2 2 3 v q v J q q\u03c9 \u23a1 \u23a4 \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5=\u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 \u23a3 \u23a6 (8) where J is the Jacobian matrix. J is a triangular matrix in the case of the parallel manipulators with decoupled motions presented in this paper. In the constraint singularity, the T2R1-type parallel manipulators with bifurcated motions have instantaneously iM=iSF=4 and (iRF)=( 1v , 2v , \u03b1\u03c9 , \u03b4\u03c9 ). The bifurcation in this constraint singularity can be used to change motion type of the moving platform. In the two distinct branches M=SF=3 and (RF)=( 1v , 2v , \u03b1\u03c9 ) or (RF)=( 1v , 2v , \u03b4\u03c9 )." + ] + }, + { + "image_filename": "designv11_92_0002029_ilt-11-2016-0283-Figure1-1.png", + "original_path": "designv11-92/openalex_figure/designv11_92_0002029_ilt-11-2016-0283-Figure1-1.png", + "caption": "Figure 1 The schematic of the air foil bearing of bump type", + "texts": [], + "surrounding_texts": [ + "k = index for search step (PSOmethod); and r = order of iterative method [Equation (4)].\nAn important aspect of tribological computing is to obtain the results in a fashion of minimum execution time. This can be achieved by using powerful computers and parallel computing, as well as efficient algorithm. Also, the selection of computing platform and high-level general-purpose programming language may affect the computing efficacy. Among the available computing tools, the use of MATLAB programming in tribological analyses is very popular, which has been implemented in many recent studies. Part of the applications of MATLAB programming in recent tribological studies is briefly reviewed and shown in the following paragraphs. An early application of MATLAB in tribological study was carried out in 2007. An efficient finite element method (FEM) was proposed and applied to solve the two-dimensional steadystate nonlinear heat conduction equation (Moreno Nicola\u2019s et al., 2007). The results are the temperature fields in hydrodynamic bearings. In that study, the FEM was coded using a MATLAB program and the execution time required was 4 min for a typical simulation. In 2013, a complex computation was conducted to solve a detailed modeling of fluid flow through a tribological interface under starved lubrication conditions (M\u00fcller et al., 2013). The execution times required were in the range of 5.42 h to 31.92 h by using MATLAB coding. For this or similar extended simulations, the use of MATLAB coding for numerical computations may not be the best choice. The development of efficient algorithms is also commonly seen in tribological studies. The goal of a recent study was to develop efficient dynamic models for a dual clearance squeeze film damper (Zhou et al., 2013). All the calculations were performed in MATLAB on a desktop computer with a 2.83- GHz quad-core processor. In a typical case, the CPU time required in solving the incompressible-fluid Reynolds equation is 7,186 s, which can be reduced to 483 s with the proposed simplified model. Similarly, a model of order reduction procedure was derived in 2015 for minimizing the computing time required to solve an elastohydrodynamic problem (Maier et al., 2015). In that study, all the calculations were performed in MATLAB using an Intel i5-2500 CPU. In a case presented, the execution time was reduced from 4,156-131 s by using the proposed scheme. In 2014, the simulative studies of tribological interfaces with partially filled gaps were conducted (M\u00fcller et al., 2014). All the computations were executed using MATLAB programming. The steady state solution was found by iteratively solving the non-stationary Navier\u2013Stokes equation until the state variables converged. A new numerical technique was developed which reduced the computational time by a factor of 1,000. As a result, the time required to compute a typical simulationwas reduced to less than 30min. Some other studies also incorporated MATLAB programming with commercial finite element programs to complete the tribological analyses (Liu et al., 2013; Fu et al., 2014; Wang et al., 2014; Zhang et al., 2016). Furthermore, the MATLAB has a user-friendly interface for simple\ncomputations (Kumar et al., 2014; Krick et al., 2014) and useful built-in functions to assist tribological analyses (Nicoletti, 2013; Liu et al., 2013; Suh and Palazzolo, 2015; Khanam et al., 2016). TheMATLAB also provides a collection of routines, such as Optimization Toolbox, which was used in a recent tribological study for a surface texture shape optimization (Shen and Khonsari, 2016). The MATLAB programming can also be used to display the computing results in real time or at the end of the execution (Cyriac et al., 2015). In contrast, the users of Fortran and C/C11 are usually postprocessing the simulation results by some standalone graphing software. On the other hand, the importance of bringing a modern or future computers equipped with multiple processor cores (MPCs) to its full computing potential for tribological component designs and/or friction, lubrication and wear analyses cannot be ignored (Chapman et al., 2007; Wang and Chang, 2012; Wang et al., 2012, 2013). Currently, MATLAB programs can invoke built-in parallel programming directives usingMPCs and/or graphics processing units (GPUs). But, the raw computing power of MATLAB codes in sequential computing (using one processor core) is disputed, particularly for computationally intensive applications. Note that the ideal performance gain that can be obtained by a parallel paradigm (computing model) is directly related to its sequential computing performance. Presently, the most popular standard of shared-memory MPC computing used in Fortran and C/C11 programming is OpenMP (Wang et al., 2015; Chan et al., 2014; Han et al., 2015). The other popular GPU-aware languages are CUDA (Wang et al., 2012, 2013) and OpenACC (Wang et al., 2013), which also support Fortran and C/C11 programming. However, the use of OpenMP/OpenACC directives or proprietary CUDA to perform parallel computing involves rewriting the legacy codes that are usually optimized for sequential computing. Also, the CUDA of the GPU computing is not compatible with standard Fortran or C/C11 language and is supported only by limited compilers. In a computer, the computing of matrix operations in MATLAB is as efficient as in using Fortran or C/C11. However, a system of linear equations obtained from a discretized partial differential equation, such as various type of compressible-fluid Reynolds equations, in tribological studies is usually solved by iterative methods (Wang et al., 2011). The use of iterative methods versus direct matrix operations (methods based on Gaussian elimination algorithm) can have the following advantages: much less memory storage requirement; easy to implement in parallel computing with very good\nspeed-up performance (Wang and Chen, 2004; Wang and Tsai, 2006; Wang et al., 2012); can handle a bearing model with recesses directly, as in modeling hydrostatic bearings; can be faster in execution if the convergence criterion is properly set (Wang et al., 2010); and in some simulations where the Reynolds boundary condition is needed, can bemet without additional treatment.\nThe use of iterative methods for solving the steady-state Reynolds equation (elliptic partial differential equation) is the\nFluid-film lubrication\nNen-Zi Wang and Hsin-Yi Chen\nIndustrial Lubrication and Tribology\nVolume 70 \u00b7 Number 6 \u00b7 2018 \u00b7 1002\u20131011\nD ow\nnl oa\nde d\nby U\nni ve\nrs ity\no f\nSu nd\ner la\nnd A\nt 0 1:\n35 1\n2 Se\npt em\nbe r\n20 18\n( PT\n)", + "basis for many fluid-film lubrication analyses. The iterative solution procedure is required in many hydrodynamic, elastohydrodynamic and thermohydrodynamic lubrication analyses. A similar iterative procedure is also required for nonlinear Reynolds equations (e.g. models for gas lubricated bearings), which are usually linearized and then solved iteratively (Heshmat et al., 1983; Wang and Chang, 1999). In the case of seeking a solution for the generalized Reynolds equation (Dowson, 1962) or its application (Wang and Seireg, 1994), the bottleneck of numerical computation is still in solving the underlying elliptic equation. To ensure a satisfactory computing performance, as well as to visualize the results obtained in bearing analysis, a study was conducted using mixed programming languages of Visual Basic and Fortran (Cui et al., 2011). However, the study was carried out in serial computing. And the computing efficiency can be improved significantly if parallel computing is applied. In this study, a performance comparison of the MATLAB and Fortran programming in sequential computing is carried out initially. The discretized Reynolds equation for modeling a simple hydrodynamic lubrication is solved by the standard successive over-relaxation (SOR) method using both the languages. As the SOR method can be easily coded with little variation between the two programming languages, the possible influence on computing efficiency due to the coding variation of the languages is minimized. Later, a cross-platform model combining MATLAB and Fortran programming is presented. The model uses the MPCs of a workstation by OpenMP directives in Fortran and the online display capability of the MATLAB is also incorporated.\nTwo inclined-surface slider models (with and without a central recess) for simulating a simple hydrodynamic lubrication are used in this study. The lubrication models are described by a non-dimensional Reynolds equation [equation (1)]. The purpose is to study the computing performance of the two programming languages, i.e. MATLAB and Fortran, in using iterative solution methods. To test the cross-platform paradigm, an optimization analysis of an air foil bearing is conducted. In the foil bearing model, the air is treated as Newtonian fluid and the variation of viscosity and density is neglected. In addition, the condition of no-slip at the bearing surface interface is also assumed. Thus, the nondimensional Reynolds equation for modeling the air bearing lubrication can be expressed as equation (2) (Wang and Chang, 2012; Wang et al., 2013; Heshmat et al., 1983):\n@ @x h 3 @p @x 1 @ @y h 3 @p @y \u00bc 6 @h @x\n(1)\n@\n@x ph 3 @p @x 1 @ @y ph 3 @p @y \u00bc 6 @ ph @x\n(2)\nIn equations (1) and (2), x \u00bc x B ; y \u00bc y B ; p \u00bc ph20 m [B ; h \u00bc h h0 , h0 = hmin for the sliders and h0 = c for the air foil bearing. In this study, equation (1) is discretized by the standard second-order accurate difference equations to form the\ndifference equation. The equation is then solved by the standard SOR method. The ranges of variables used in the slider models are: 0 x 1, 0 y 1 and h x\u00f0 \u00de \u00bc 11 1 x\u00f0 \u00de, and the grid sizes are the same in the x- and y-directions (Dx \u00bc Dy). For the recessed slider, the central recess occupies the range of 0:4 x 0:6 and 0:4 y 0:6. In the simulations, the film pressure at the boundaries for the both sliders is set to zero and the pressure in the recess is also assumed to be zero. In the foil bearing, the flexible foil surface is supported by a bump foil which is lubricated by air (Figures 1 and 2). The foil and the supporting bumps are simulated by an elastic model (Heshmat et al., 1983) to determine the foil deformation in the bearing with a known eccentricity ratio. Table I lists the values of the parameters and operating conditions of the foil bearings, as well as the two design variables and their range. In this study, equation (2) is linearized by Newton\u2019s method and solved iteratively (Wang andChang, 2012).\nIterative methods and relaxation factors\nIn this study, the point SOR method with natural ordering is applied to solve the discretized equations (1). The SOR method is carried out in sequential computing by either MATLAB or Fortran. The SOR method used to solve equation (1) contains a two-dimensional matrix of elements, where each element in thematrix represents a film pressure and is updated by taking its average along with the four weighted neighboring elements in the grid.\nFluid-film lubrication\nNen-Zi Wang and Hsin-Yi Chen\nIndustrial Lubrication and Tribology\nVolume 70 \u00b7 Number 6 \u00b7 2018 \u00b7 1002\u20131011\nD ow\nnl oa\nde d\nby U\nni ve\nrs ity\no f\nSu nd\ner la\nnd A\nt 0 1:\n35 1\n2 Se\npt em\nbe r\n20 18\n( PT\n)", + "The Newton\u2019s method is applied to linearize equation (2) and solved by the red-black SOR (RBSOR) method. During the iterative procedure, the computation for the film pressure is arranged in red-black ordering. In the red-black ordering, the grid is divided into \u201cred\u201d points and \u201cblack\u201d points, which are interweaved. The black points are computed using four neighboring red points and the red points are computed using four neighboring black points. The computation is organized in two phases that are repeated until the values converge. All black and red points can be computed simultaneously in a sequential order. It has been shown that the execution of RBSOR for the linearized Reynolds equation is faster than using the standard point SORmethod either in serial or parallel computing (Wang andChang, 2012). In this study, equation (3) is used to calculate the optimal relaxation factor (v opt) for the SOR method. The constant (a) in equation (3) is determined empirically by running several cases of various grid sizes. The dimensionless grid size (Dx) is a reciprocal of the grid number used and Dy \u00bc Dx . While v opt is the optimal asymptotic relaxation factor, it is not necessarily a good initial choice (Wang et al., 2012). In this study, the a in equation (3) is 7.633 or 2.961 for the slider with or without the central recess, respectively:\nv opt \u00bc 2 11aDx\n(3)\nIn a sequential computation, both the MATLAB and Fortran programs are coded in a similar way, except some variations due to the language syntax difference. In the coding for the recessed slider, the calculation of the film pressure within the recessed region (which is a constant) is skipped by a logic ifstatement in the iterative nested-loops. Therefore, the film pressure in the recess remains zero, which is initialized at the beginning of the iterative computation of the film pressure. A similar operation may be required in an iterative solution method for tribological analyses, whichmay adversely influence the execution time in nested-loops.\nStopping criterion of iterative method\nAccording to Wang et al. (2010), a reasonable stopping criterion of iterative methods for solving Reynolds equation is to terminate the iterative procedure when the L2 norm of\nresidual, Dx2 Xn i\u00bc0 Xn j\u00bc0 Dpi;j 2 !1 2 , in iterative computation is smaller than the truncation error (CDxr) of the difference method used. Thus, the iterative procedure in this study is terminated when the stopping criterion [equation (4)] is met.\nDx2 Xn i\u00bc0 Xn j\u00bc0 Dpi;j 2 !1 2 < CDxr (4)\nwhere C \u00bc 2r jp1 p2j Dx\u00f0 \u00der 2r 1\u00f0 \u00de or 22r jp2 p3j Dx\u00f0 \u00der 2r 1\u00f0 \u00de, r \u00bc log2 p1 p2 p2 p3 , Dpi;j is the\nresidual of p at grid (i, j), p1 \u00bc pa Dx\u00f0 \u00de, p2 \u00bc pa Dx 2 and\np3 \u00bc pa Dx 22 . Both C and r are determined from three coarse gridwork calculations. In this study, the truncation error is a function of the grid size and C = 0.065713 and r = 1.99818 for both the sliders. In the case of air foil bearing, the RBSOR method is used. The gridwork used is 378 126, the corresponding relaxation factor is 1.9378 and the truncation error is 1.36 10 6.\nToday\u2019s operating systems are very complex and the execution time of a program is affected more or less by some background activities. Instead of averaging the execution time from several runs, the execution time obtained and used in this study is by running a case three times and recording the one with the shortest elapsed time. This should better reflect the actual time required in an ideal computing environment. In this study, the effectiveness of a programming language on computing is assessed by using the language-dependent performance indicator (g) defined in equation (5). This indicator removes the effect on execution time due to the differences of total number of iterations (N) and the size of gridwork (n2) in various computations. As a result, the indicator provides the normalized execution time in one iteration of a grid. For a given machine, this indicator can be viewed as a direct measure of the computing capability of a programming language being used.\ng \u00bc tS N n2\n(5)\nThe speedup factor (S) defined in equation (6) is a ratio of the sequential computing time (ts) to the elapsed time (tp) required in parallel computing. The amount of speedup obtained in a computational task is influenced by the numerical methods, programming platform and computing machine. The ideal speedup of a parallel computation in a system can approach the number of available processor cores in the computing:\nS \u00bc ts tp\n(6)\nParticle swarm optimization method\nIn this study, a standard particle swarm optimization (PSO) computational algorithm is applied (Kennedy andEberhart, 1995)\nFluid-film lubrication\nNen-Zi Wang and Hsin-Yi Chen\nIndustrial Lubrication and Tribology\nVolume 70 \u00b7 Number 6 \u00b7 2018 \u00b7 1002\u20131011\nD ow\nnl oa\nde d\nby U\nni ve\nrs ity\no f\nSu nd\ner la\nnd A\nt 0 1:\n35 1\n2 Se\npt em\nbe r\n20 18\n( PT\n)" + ] + } +] \ No newline at end of file